ML19316A040

Core Flooding Nozzle Mechanical Design.

ML19316A040
Person / Time
Site:	Oconee
Issue date:	01/15/1973
From:	BABCOCK & WILCOX CO.
To:
Shared Package
ML19316A038	List: ML19316A037 ML19316A039 ML19316A040
References
NUDOCS 7911210602
Download: ML19316A040 (65)

Text

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• ATTACIDfENT I

_ CORE FLOODING NOZZLE MECHANICAL DESIGN l

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. LIST OF FIGURES 1 Existing Core Flooding Nozzle Sleeve j 2 Oconee I Core Flooding Nozzle Insert

. 3 Core Flooding Nozzle Af ter Removing Existing Sleeve l 4 Core Flooding Nozzle Weld Preparation Prior to Inserting Modified i

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- Insert s

i 5 Core Flooding Nozzle Insert Installed in Core Flooding Nozzle 1

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n The core flooding nozzle modification for the Oconee I nuclear plant is basically opening from a variable 0.72 ft2diameter thermal sleeve which reduces the nozzle to 0.44 ft2 rication, and field installation of the sleeve.This section reviews the design, fab-The

1. existing core flooding nozzle and thermal sleeve are shown in Figure The existing sleeve will be removed to enable installation of the modified sleeve.

The restrictor is fabricated in two scages as shown va Figure 2. Because of schedule it was necessaryrestraints to userelatedavailable to installing raterial. this modification at Oconee, Thus, Type 304 stainless steel pipe was used and weld overlay was deposited to meet _the required 9" I.D.

The pipe and weld overlay were U.T. examined before machining with P.T.

examination af ter final machining. These examinations were performed in accordance with ASME Section III.

Removal of the existing sleeve is accomplished by grinding the weld buttons which hold it in3).place, and performing a P.T. examination on the ground areas (Figure

. The nozzle is then ready for installing the restrictor.

ring as shown in Figure 4.by welding and machining the weld buttons and This will be accomplished The restirctor is inserted and a full penetra-tion weld with permanent backing ring is made in accordance with ASME Section III (Figure 5).

ity weld. A progressive P.T. is performed to insure a qual-at The weld buttons center the restrictor. A drain hole is drilled the bottom strictor to prevent of the crud weld to allow a small flow of water behind the re-buildup. .

l The restrictor and attachment veld (Figure 5) are evaluated in accordance with ASME Section III. The significant transients which affect the re-strictor and weld are reactor coolant system heatup and cooldown including the core flooding system periodic test transient and decay heat removal initiation.

All transients are considered as normal operating conditions and factor. are considered in determining thermal stresses and the fatigue usage on the weld per ASMEanalysis The fatigue Section III.includes a strength reduction factor of two The weld has also been designed to with-stand poi the faulted condition where a differential pressure of up to 2250 ,

may occur because of a core flooding line LOCA. A dynamic magnifica- )

tion factor of two was applied to the pressure to account for instantaneous application. l yields a t4afety Basedmargin onofthese 1.4. criteria, the average shear stress in the weld These assumptions and safety marg'n i are sufficient to insure the structural integrity of the ndzzle, restrictor, and weld for all operating and faulted conditions.

During the core flooding transient, the maximum Ap across the nozzle is

! expected to be approximately 200 psi. This is a factor of greater than 20 less than the design loading assumptions. 'Therefore, it is not consi-dered credible that the restrictor retaining weld would fail during core flooding tank discharge.

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During operation of the decay heat system, the Ap loads on the restrictor are insignificant.

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l FIGURE 1 EXISTING CORE FLOODING N0ZZLE SLEEVE s s 1

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FIGURE 2 OCONEE I CORE FLOODING N0ZZLE INSERT g[R

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O FIGURE 4 CORE FLOODING N0ZZLE WELD PREPARATION PRIOR TO INSERTING MODIFIED INSERT s ,

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T / - _- .. _ . , 2#_- ...- .m.1_ . - _.. _,, .-__m __ m ._21__. ._ .-- m . . . . .__ --- j. I ^ f a a 1 I 1 i i i ATTACHMENT II t b { ANALYSIS OF A CORE FLOODING LINE BREAK LOSS OF COOIANT ACCIDENT FOR THE OCONEE 1 REACTOR WITH INSERT IN CFT NOZZLE l, 1 I i l } l l l J t E' .s 4 e --ne ,c-. y - , --, - v -w,- ,,- e , - -,-ng,- , w,,,-=-w,- - - - . - + = --~ e 1 . LIST OF FIGURES 3-1. CRAFT Evaluation Model 3-2. Sensitivity of Inner Vessel Liquid Volume to Noding 4-1. Core Pressure Versus Time 4-2. Core Power Versus Time 4-3. Inner Vessel Fluid Volumes Versus Time 4-4. Vent Valve Flow Versus Time 4-5. Leak Flow Versus Time 4-6. Water Height in Core and Downcomer Versus Time 4-7. Fluid Velocities From the Downcomer to the Lower Head During LFT Injection 4-8. Axial Power Shape for Case 1 1 4-9. Case 1, Level 5 - Sink Temperature Versus Time 4-10. Case 1, Level 5 - Heat Transfer Coefficient Versus Time 4-11. Case 1, Level 5 - Cladding Temperature Versus Time 4-12. Case 1, Level 9 - Sink Temperature Versus Time 4-13. Case 1, Level 9 - Heat Transfer Coefficient Versus Time 4-14. i Case 1, Level 9 - Cladding Temperature Versus Time l 4-15. i Case 1, Level 10 - 91nk Temperature Versus Time 4-16. Case 1, Level 10 - Heat Transfer Coefficient Versus Time ' 4-17.  ! i Case 1, Level 10 - Cladding Temperature Versus Time I I 4-18. Axial Power Shape for Case 2 }4-19. Case 2, Level 7 - Sink Temperature Versus Time 4-20. i Case 2, Level 7 - Heat Transfer Coefficient Versus Time  ! 4-21. Case 2, Level 7 - Cladding Temperature Versus Time 4-22. Case 2, Level 9 - Sink Temperature Versus Time 4-23. Case 2, Level 9 - Heat Transfer Coefficient Versus Time 4-24. Case 2, Level 9 - Cladding Temperature Versus Time I t II-i I ' i - - ~ LIST OF FIGURES (Continued) 4-25. Case 2 Level 10 - Sink Temperature Versus Time 4-26. Case 2, Level 10 - Heat Transfer Coefficient Versus Time 4-27. Case 2, Level 10 - Cladding Temperature Versus Time 4-28. Axial Power Shape for Case 3 , 4-29. Case 3, Level 9 - Sink Temperature Versus Time 4-30. Case 3, Level 9 - Heat Transfer Coefficient Versus Time 4-31. Case 3, Level 9 - Cladding Temperature Versus Time 4-32. Case 3. Level 10 - Sink Temperature Versus Time 4-33. Case 3, Level 10 - Heat Transfer Coefficient Versus Time 4-34. Case 3. Level 10 - Cladding Temperature Versus Time i 3 II-ii

1. INTRODUCTION The scope of this analysis is a guillotine break of the core flood tank line between the reactor vessel nozzle and the first check valve. Flow out of the reactor vessel is limited to an ef fective area of 0.44 f t2 due to a flow limiting insert.

Since the leak ayea is less than 0.5 ft2, the B&W small leak evaluation medel, BAW-10052 , "Multinode Analysis of Small Breaks for B&W's 2568-MWt Nuclear Plants", is used. Consideration is given to three different axial power distributions. The following assumptions are made for conditions and system responses during the accident:

1. The reactor is operating at 102% of the steady-state power level of 2568-MWt.

2. A single failure is assu,med in addition to the CFL break. The worst single failure results in an injection flow from only one high pressure injection pump and the second core flooding tank.

3. The leak occurs instantaneously, and a discharge coefficient of 1.0 is used for the entire analysis.

4. The reactor trips on low pressure at 2050 psig.

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2.

SUMMARY

AND CONCLUSIONS The maximum cladding temperature for this analysis is a function of the axial power shape. For the three axial power shapes analyzed, the maxi-mum cladding temperature is 1172 F. A peak temperature of this magnitude or less presents swelling; no potential for either metal-water reaction or cladding therefore, the core geometry remains unchanged and amenable to cooling. This analysis indicates that adequate core cooling is maintained.

The operator may initiate additional injection to refill the vessel in about one-half hour. Even without this additional long term cooling estab-lished, however, all conditions of the AEC Interim Acceptance Criteria are met.

i The peak temperatures for the three different axial power shapes analyzed in this report are as follows:

Elevation of Elevation of Peak Peak Power Peak from Cladding Temperature Cladding

' Bottom of Core from Bottom of Core Temperature ft ft F 1

5.5 5.5

! 663 7.8 11.4 666 10.6 11.4 1172 I

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3. METHOD OF ANALYSIS The method of analysis used to determine the cladding temperature res ponse is the same as described in BAW-10052.

policy statement because, although the ThisCFT is consistent with the interim line cross-sectional area is, 0 72 f the flow from the reactor vessel is limited in the vessel which has a cross-sectional area of 0.44 ft2, z e by annoz insertl The basic assumptions of the CRAFT model, used for core hydrodynamics, scheme is somewhat different due to the nature The noding of the break scheme and legend are shown in Figure 3-1.

A THETA-1B controlled region.model is used to determine the heat transfer in the flow sufficient detail the three axial power shapes that are . The analyzedThe Quenchismodel, regime not flowas described in BAW-10052, is used when the heat transfer controlled.

When the flow in the core has decayed sufficiently, the core is in a rela -

tively quiescent condition in which the lower portion is covered by a two-cooled at a flow rate consistent with the boil-off rate in t This steam icant temperaturecooling results in reduced heat transfer coefficientsnasignif-transients. d Therefore, it is very important to properly and conservatively determine the height of the two phase mixture Steam .

production as calculated by CRAFT and the Redfield-Murphy bubble4 rise model is used to determine the two-phase mixture height.

limiting for thermal void fraction in the inner vessel mixtureConservatism the analysis. to a maximum is applied value ofby0 5 .

mixture levels in the inner vessel.The void fraction as calculated by CRAFT results in higher By limiting the void fraction to 0.5, more of the core in the uoner region.is uncovered, which results in less steam flow and more superheating The CRAFT schemes. model was examined to determine its sensitivity to various noding Figure 3-2 shows the results of three noding approaches in terms of inner vessel liquid volume and confirms our conservatism in using a two node inner vessel model because it eliminates the " pancaking"effect.

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4. RESULTS OF ANALYSIS The CRAFT run was made with a single core flow path and is applicable to any of the three axial power shapes examined. Figures 4-1 and 4-2 show the core pressure and power history for this accident. Figure 4-3 depicts the inner vessel fluid volumes as used for the heat-up calculations. The vent valves are located above the top of the active region; therefore, the vent valve flow shown in Figure 4-4 will appear to be inconsistent with the mixture height shown in Figure 4-3. In section 3, it was explained that a conser-vative limit was placed on the core void fraction. This effect is visible here because CRAFT shows a mixture height in the arca of the vent valves while the nixture height which is used in the thermal analysis is much lower. Mixture flowing out of the vent valves is conservative because it removes liquid water from the core region.

The Icak flow as a function of time is shown in Figure 4-5. The core and downcomer water heights are shown in Figure 4-6, while Figure 4-7 shows that calculated fluid velocities between the core and downcomer are not of sufficient magnitude to consider entrainment of the CFT water.

The 10 axial region THETA model was used until 300 seconds for the three axial profiles shown in Figures 4-8, 4-18, and 4-28. The Quench code was then used to analyze the cladding thermal response for the remainder of the transient.

The center-peaking power shape used in analyzing Case 1 is shown in Figure 4-8. The upper portion of the core is uncovered during the transient; but the low power obtained in this region does not result in significant clad-ding temperature increases. Therefore, the center region of the core, axial level 5, remains the hottest portion of the core with a peak cladding temperature of 663 F. Figures 4-9 through 4-11 show data related to axial level 5 and Figures 4-12 through 4-14 and 4-15 through 4-17 are relevant to axial levels 9 and 10 respectively.

Outlet power peaks are important to this enalysis because higher power re-gions will be uncovered and the amount of superheat in the cooling steam will be higher. Case 2, Figure 4-18, is typical of an outlet peak exper-ienced in the core. The maximum power peak occurs in axial level 7. It g

is never uncovered by the two-phase mixture, and the cladding temperature remains low. Figures 4-19 through 4-21 provide information on this level.

g Axial level 9 is uncovered for only a short period of time between 740 and I 770 seconds and does not undergo a large cladding temperature rise. Figures 4-22 through 4-24 apply to this level. The highest peak cladding tempera-ture is achieved at level 10 which is uncovered from approximately 1300 seconds to 2000 seconds and reaches a peak cladding temperature of 656 F.

Information for this Icvel is shown in Figures 4-25 through 4-27. The cladding temperature is declining slightly at the end of the analysis.

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To see the effect of an outlet peak, Case 3 which is shown on Figure 4-28 was cho.sen for examination. This power shape represents one of the most adverse tilts toward the exit of the core. This shape is used for design purposes and while it is not an expected or normal operating condition, this shape is allowable for operation for the last few days of core life.

In the heatup calculation, the power peak occurs at axial level 9. The temperature at the location of peak power, level 9, does rise as it is uncovered, but the recovering of the level keeps this from being the worst loca tion. Figures 4-29 through 4-31 provide information at this location.

The peak cladding temperature, 1172 F, occurs at level 10 because of the extended steam cooling from 1300 seconds on. The cladding temperature, Figure 4-34, is high because of the low heat transfer coefficient, Figure 4-3% and the degree of superheat, Figure 4-32. Both effects are caused by the shift of power to the upper regions of the core, i

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5. HIGH PRESSURE INJECTION FLOW i

The analysis of the core flooding line break shows that the core can be cooled using only one core flooding tank and one high pressure injection (HPI) pump. The HPI flow used in the analysis was based on test results from Oconee 1. A least-squares regression analysis of the data was performed which showed a flow of 352 gpm and 353 gpm in each of two strings at a RCS pressure of 1500 psig. Considering an instrumentation error of i 1%

and the relative error in the regression fit, the flow at 1500 psig was reduced to 340 gpm for use in the analysis. Similarly at 600 psia, the least-squares fit resulted in a flow of 457 gpm which was reduced to 440 gpm for use in the analysis. Using these points, together with the tested pump head capacity curve, the high pressure injection over the full pressure range was established.

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6. LONG TERM ECCS OPERATION The preceding analyses were based on ECCS capability from one HPI pump and one core flooding tank. This condition assumed that the active LPI pump was lined up to pump to the core flood line that had the break and the other LPI pump was inoperative by the criteria of a single active failure.

Increased long term safety margin can be obtained by operator action to initiate low pressure injection through the unbroken core flood line. This action can easily be taken during the first 30 to 60 minutes after the CFT line break. The operator will open control room operated cross connect valves at the LPI purp discharge and check flow indicators in each of LPI lines to determine that some flow is going through each line. Equalization of flow in the lines can be accomplished from the control room by posi-tioning of control valves in each LPI line. When flow is equalized through each line the LPI flow into the reactor vessel will be at least 1500 gpm with one pump operating and 3000 gpm if both pumps are operating.

The Oconee station operating procedures will be changed as follows:

1. Prior to switching suction on the ECCS and RB Spray pumps from the BWST to the RB sump or before shutting off all HPI pumps, check LPI flow indication and LPI pump operation to assure flow into the reactor vessel. This requires flow indication in each line since it is not known which line has the break.

2. If only one pump is running the operator should take the follow-ing actions.

a. Attempt to start idle LPI pump. Failure to start may be ES actuation failure. Operator can operate valves and start pumps by remote manual control from control room.

b. If pump (LP-PIC) is available,' place in operation on the LPI string where pump is not running by opening valves in suction and discharge crossover lines and starting pump LP-Plc from the control room. Observe flow indication in the LPI line.

This action produces 3000 gpm through each LPI line.

c. If operator cannot start either of the two LPI pumps (s teps a & b), perform the following steps to achieve flow into the reactor vessel from the one active LPI pump:

Open discharge crossover valves to get LPI flow into each of the LPI lines.

Monitor LPI flow indication to assure flow through each line.

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Adjust the throttle valve in each LPI line until a flow balance is achieved. This will give approximately 1500 gpm through each line.

All valves also have local handwheels that can be manually actuated.

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7. COMPARISON OF EVALUATION MODEL WITH APPENDIX A, PART 4 0F THE AEC INTERIM ACCEPTANCE CRITERIA FOR EMERGENCY CORE COOLING SYSTEMS Although Apgendix A, Part 4 is strictly appropriate only for breaks larger than 0.5 ft , a check list comparison of that evaluation model to the one used in this report may be of convenience to the reader. Evaluation and explanations are provided on a point by point basis with a subdivision con-sistent with Appendix A, Part 4.

The first paragraph of Appendix A, Part 4 lists several reports written by B&W which document techniques to be applied in the large break evaluation mod e l . These reports are appropriate as follows:

1. CRAFT - This report and the code described are used f or the CFT line break.

2. REFLOOD - This code is used for the purpose of calculating the refilling of a vessel once that vessel has reached end of blow-down. As that situation does not occur for the CFT line break, the code and its report do not apply and are not used. ,

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3. THETA 1-B - This report and the code described are used for the CFT line break.

4. BAW-10034 - This report is written for large breaks.

Appendix A, Part 4 goes on to list specific instructions for the large break evaluation model.

l 1.1 Core and System Noding I 1

1.1.1 Only one core node has been used in the CFT line break analysis.

( 1.1.2 The Theta model used during the flow controlled heat transfer l regime had 6 fuel nodes, 2 clad nodes, and 10 axial levels. ,

After the flow controlled heat transfer regime, after 300 l seconds, the Quench code is used. This code has 1 fuel and l 1 clad node and must be applied individually at separate '

axial levels.

1.2 Pump Model This model is the same used ir. BAW-10052 and is discussed in that report. It is different from the model used in large break analysis though both models are consistent with the Appendix A, Part 4 guideline .

1.3 Break Characteristics This statement does not apply to a specific break like the CFT line l break. '

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r 1.4 bischarge Coefficient As suggested by Appendix A, Part 4, a discharge coefficient of 1.0 has been used.

1.5 Decay Heat The decay heat cury suggested in Appendix A, Part 4 was used in the CFT line break analysis.

1.6 Time to Departure from Nucleate Boilinn (DNB)

This was done as suggested in the large break evaluation model.

1.7 Film Boiling Heat Transfer This was done as suggested by the large break evaluation model except that for pool film boiling used in the QUENCH code, the Morgan corre-lation was used.

1.8 Metal-Water Reaction Rate This was done as suggested by the large break evaluation model. Tempera-tures for this accident, however, prohibit any significant metal-water reaction.

1.9 Core Flow Rate This was done as suggested by the large break evaluation model while flow was controlling the heat transfer.

1.10 Enthalpy and Pressure This was done as suggested by the large break evaluation model.

1.11 Core Flooding Tank Bypass As downcomer steam flows were insufficient to cause entrainment of core flooding tank water, this bypass assumption was not imposed on the CFT line break analysis, i AppendixA,Part4thenproceedstodescribetheevaluationmodelforfhe reflood portion of the large break. As there is no classic reflood portion for the CFT line break, this section does not apply to the analysis.

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4 LIST OF REFERENCES 1.

C.E. Parks, B.M. Dunn, and R.C. Jones, Multinode Analysis of Small Breaks for B&W's 2568-MWt Nuclear Plants, RAW-10052, Babcock and Wilcox, Lynchburg, Virginia, September,1972.

2. c, GRAFT - Des'ription of Model for Equilibrium LOCA Analysis Program, BAW-10030, Babcock and Wilcox, Lynchburg, Virginia, October, 1971.

3.

C.J. Ilocever and T.W. Wineinger, THETA 1-B - A Computer Code for Nuclear Reactor February,Core 1971.Thermal Analysis, TN - 1445. Idaho Nuclear Corp.,

4.

J.A. Redfield and J.H.. Murphy, Void Fraction and Residual Water Predictions During Loss of Coolant, WAPD-T-2155, Westinghouse, BAPL, September 1968.

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1. LIST OF REFERENCES C.E. Parks, B.M. Dunn Lynchburg, VirginiaBreaks-

, and R.C. Jones,for B&W's Multinode A 2568 MW 2 t Nuclear

, September, Plants, BAW 1 1972 nalysis of Small CRAFT - Desc'ription

, - 0052, Babcock and Wilc ox,

3. BAW-10030, Babcock and'Wof Model for Eauilibrium L C.J. Hocever Reactor ilcox, Lynchburg, Vir iOCA Analysis Program February, Core Thermaland T.W. Wineinger, g nia, October,1971 THETA 1 1971.

1 4 Analysis, IN - 1445-B - A Computer Code f

, Idaho Nuclear Corp., or Nuclear Predictions During .

LoJ.A. Redfield and J H

.. Murphy, Void Fraction ss BAPL, September 1968 .

of Coolant, WAPD-T-2155and Residual Water

, Westinghouse, 1

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FIGURE 3-2 SENSITIVITY OF INNER VESSEL LIQUID VOLUME TO NODING .

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FIGURE 4-6 WATER HEIGHT IN CORE AND DOWNCOMER VERSUS TIME 36 32 -

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FIGURE 4-7 FLUID VELOCITIES FROM THE DOWNCOMER TO nic LOWER HEAD DURING CFT INJECTION 6

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FIGURE 4-8 AXIAL POWER SHAPE FOR CASE 1 2.0 .

1.8 1.6 1.4 1.2 g Ax al Peak = 1.786 1'0 E Radial Peak = 1.7157 2 0.8 - - Axi al Power Sha'pe

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O

3 10 0 -

m E -

102 m

101 0 400 l

800 1200 1600 2000 Time, s I

i 9

0 0

0

~ 2 0

0 8

1 E

M I

T S 0 U 0 S 6 R 1 E

V E

R U 0 T 0 A 4 1

R E

P M

E T

0 G 0 2

N 1 I

D D

A s L

C ,

0 e

- 0 0 i m

1 T 5

L E

V E

L 0 0

8 1

E S

A C

0 0

6 1

1 4

0 E 0 R 4 U

G I _

I F

- 0 0

2

\

0 0 0 0 0 0 0 0 0

_ 0 0 0

_ 4 2 0 0 0 0 0

1 4 8 6 1 4 2 w,i3

. ; = 3".a L

FIGURE 4-12 CASE 1, LEVEL 9 - SINK TEMPERATURE VERSUS TIME 650 9 600 550 500 E

E 450 '

O t

a 400 h .

350

~

300 _

0 200 400 600 800 1000 1200 H00 IW M E Time, s

FIGURE 4-13 CASE 1, LEVEL 9 - HEAT TRANSFER COEFFICIENT VERSUS TIME 5

10

~

~

.104 t

"I T -

E h 103 m -

I E -

G I -

I -

0 102 0 -

g -

t N

E -

W 10I -

100 0 200 400 600 800 1000 1200 1400 1600 1800 2000 Time, s

I FIGURE 4-14 CASE 1, LEVEL 9 - CLADDING TEMPERATURE VERSUS TIME 1400 1200 1000 t."

800 2

O g 600 i 400

[ k 230 0

0 200 400 600 800 1000 1200 I400 1600 1800 2000 Time, s

l ll 0

0 0

2 0

0 8

1 E

Df T

0 S 0 U 6 1

S R

E V

E R

U 0 0

T 4 A 1 R

E P .

M E

T 0

0 K 2 N 1 I

S 0 0 s

1 0

' 0 e, L 1 m E i T

V E

L s

0 1 0 8

E -

S A

5 1

C J 0 0

6 4

E R

U G

I F

M 0

0 4

0 0

2 0 0 0 0 0 0 0 0 4

1

. 2 1

0 0 0 8

0 8

0 0

0 1 2 7 i 5 " .E~ sG i

FIGURE 4-16 CASE 1, LEVEL 10 - HEAT TRANSFER COEFFICIENT VERSUS TIME 105 W

m M

4 10 d

e. ms 5

103 m .

E -

E

[ 10 2 -

- s N E

101 - -

W e

W 100 , ,

0 400 800 1200 1600 2000 Time, s

FIGURE 4-17 CASE 1, LEVEL 10 - CLADDING TEMPERATURE VERSUS TIME 1200 .

i000 800 '

?

5 600 i m

i -

f i 400 um /

/

~-

200 0

0 200 400 600 800 1000 1200 1400 1600 1800 2000 Time, s

FIGURE 4-18 AXIAL POWER SHAPE FOR CASE 2 1.6 1.4 1.2

\

__ -- a 1.0

[

2 _ __ _. - - - -

• 0.8 Axial Peak = 1.533 Radial Peak = 1.665 Axial Power 0.6 Sh ap e

- - -- Power Di stri bu-tion used in THETA 0.4 _

0.2 1 2 3 4 5 6 7 8 9 10 Atlat Segment 6

_ -_._s a-  : - s FIGURE 4-19 CASE 2, LEVEL 7 - SINK TEMPERATURE VERSUS TIME 650 .

600 550 i

3 500 t

.I 3 458 '

400 3

350 3

~

300 0 200 400 600 800 1000 1200 1400 1600 1800 2000 Time, s

l l

i l

1 FIGURE 4-20 CASE 2, LEVEL 7 - HEAT TRANSFER COEFFICIENT VERSUS TIME 105 w

h me 4

10 m

b e W

R m

103

~

R~

r -

2 -

S3 0 102 the . seu E

\ ~ _

I -

101

_ Y

=

h 0

10 400 800 1200 1600 200,0 Time, s I

FIGURE 4-21 CASE 2 LEVEL 7 - CLADDING TEMPERATURE VERSUS TIME

• 1200 1000 Y

a j 800

.E .

g 600 G

400 -- -

200 0 200 400 600 800 1000 1200 W #

Time. 5 a

FIGURE 4-22 CASE 2, LEVEL 9 - SINK TEMPERATURE VERSUS TI!E 650 600 500 t

b 5 450 L E

a G -

400 350

'A 300 O 200 400 600 800 1000 1200 1400 1600 1800 2000 Time, s

FIGURE 4-23 CASE 2, LEVEL 9 - HEAT TRANSFER COEFFICIENT VERSUS TIME 5

10 M

m e

4 10

~ .

$ 103

=

J -

E o .

5 o

102 5 . _-

=

~

~

E 101 -

m W

i

~

100 i e i i 0 400 800 1200 1600 2000 Time, s

FIGURE 4-24 CASE 2, LEVEL 9 - CLADDING TEMPERATURE VERSUS TIME 1200 1000 m

J 800 600 A

!! J f\ t 400 200 0 200 400 600 800 1000 1200 1400 1600 1800 2000 Time, s

FIGURE 4-25 CASE 2, LEVEL 10 - SINK TElfPEUATURE VERSUS TIFE 650 .

600 500 d -

E

=

@ 450 '

3 a

400 T -

350 --

300 0 200 400 600 800 1000 1200 1400 1600 1800 2000 Time,s

FIGURE 4-26 CASE 2, LEVEL 10 - IIEAT TRANSFER COEFFICIENT VERSUS TIME 0

10 m

W 4

10

~

? L

~

f 103 b ,

~

E G

5 10 2

h N g -

E 101 "

O e

i

~

l 0 ,

10 , , ,

i 0 200 400 600 800 1000 1200 1400 1600 1800 2000 Time, s i

FIGURE 4-27 CASE 2, LEVEL 10 - CLADDING TEMPERATURE VERSUS TIME 1400 ,

1200 1000 t

,,- 000 E

e s00 L /

,,, VJdA 200 0

0 200 400 800 000 1000 1200 I400 1600 1200 2000 Time, s

t FIGURE 4-28 AXIAL FONZR SHAPE FOR CASE 3 1.6 1.4

[ i Antal Peak = 1.466 Radial Peak = 1.432 1.2 Axial Power shape F

~-- --

- --- Power Di stri-bution used

% - -. in THETA g 1.0 f '

."_ /

0.8 - - - - _ _

0.6

~

l 0.4 e 1 2 3 4 5 6 7 d 9 10 Atlal Segeert g

0 0

0 2

0 0

9 E

M I 0 T 0 6

1 S

U S

R E

V E 0 0

R 4 U 1 T

A R

E P

M E 0 0

T 2 1

K N

I S

0 0 s 9 0 1

e, L m E i V T E

L 0

3 w 0 8

E S

A C

0 0

9 , 6 2

4 E

R U 0 G 0 4

I F _

_ ' 0 0

2 I

0 0 0 0 5 0 0 0 0 0 5 0 0 6 6 5 5 0 5 0 5 4 4 3

. 3 7,:B* .E.* i G

FIGURE 4-30 CASE 3, LEVEL 9 - HEAT TRANSFER COEFFICIENT VERSUS TIME 105 I

• 4 10 o7 \ ~~

3 hs 10

W~

Z -

2 2 -

C -

3 102 t  % -

\

O -

c -  %

2 -

10I

~

- i

~

_ l 100 0 400 800 1200 1600 2000 Time, s

FIGURE 4-31 CASE 3, LEVEL 9 CUDDING TEMPERATURE VERSUS TIME I400 - - - . -

1200 1000 Y

f= 800 as 800 k - I k f -

1  %

=

G 400 200 0

0 100 400 600 800 1000 1200 1400 1600 1800 2000 Time, s

FIGURE 4-32 CASE 3, LEVEL 10 - SINK TDIPERATURE VERSUS TIME .

1400 --

1200 1000 p 800 B

k h 800 I

m 400 200 0

0 200 400 600 800 1000 1200 1400 1600 1800 2000 Time. s

l l

(

l

[

l FIGURE 4-33 CASE 3, LEVEL 10 - IIEAT TRANSFER COEFFICIENT VERSUS TIME 105 m

D 4

10 M

P a 103 5 [ P 5

a S 102 s

N t

a -

t -

E 1 -

10 -

=

iOO ,

I 0 400 800 1200 1600 2000 Time, s

E 8

E H

E P

m o w S w -

C4 w

w

& a

1. N a f

w w

a e 8 w

z -

o o e 5U J

E i

) a C 8

O -

M A

w w o A o

/ "

m w

m a

r 8

o m

5 I 1

4 M E o

H N

8 m

s o

8 E o a g 8 8 E * ~

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Previous analyses have shown that a marginal amount of water is maintained in the core following a core flooding tank line break. This is due to a partial degrading of the ECC System as a natural consequence of the acci-dent. In order to improve the safety of the design, a flow limiting insert is to be placed in the core flooding tank nozzle to control the accident.

By limiting the violence of the blowdown, less reactor coolant system water will be ejected from the primary system. Thus, more water will be maintained in the core during and af ter blowdown.

The acceptability of the insert is dependent on three points. First, it must be shown that it can be built and installed to function as designed.

This has been addressed in Attachment 1. Second, it is necessary to show that the insert as designed will produce the desired results. This has been addressed in Attachment II. Finally, it is necessary to show that the insert will not jeopardize the perf orme. ace of the ECC System during other accidents. This is addressed in this attachment.

By standard methods of analysis, a i-factor has been determined for the insert. The value of this factor is 0.2 based upon an area of 0.7213 ft 2 and is used to solve for the pressure loss op in the following equation:

klW [W_

3p , 288 pg A where W = flow rate, Lbm/sec p = density, Lbm/f t g = gravitional constant, Lbf e A = Area, ft The k-factor for the coce flooding line resistance used by B&W in the evaluation of LOCA's presented in BAW-10034 was 6.3. This value is typical of all plants of this Totype. Theproppsedinsertwouldincrease show that such an inc ease is acceptable, the resistance by only 3%.

an analysis of the worst case large break, an 8.5 f t cold leg split, has been carried out for two different k-f actors. The first value, k = 8.3, has an increase (6.3 + 2.0) an order of magnitude higher than the proposed insert. This was chosen before the insert had been designed in order to bound the result. The second value, k = 5.5, is based on an experimental measurement of the line k-factor without the insert, k =

4.8, plus a conservative evaluation of the insert effect, k insert =

0.7. With a more concrete design and a better evaluation of the effect of the insert, we expect the actual line resistance to be k = 4.8, plus k insert = 0.2 or k = 5.0. #

III-l

t The results of the two different k-factors are shown in Figures 1 and 2.

Figure 1 shows the core flooding tank injection rate for two tanks. Fig-ure 2 shows the resulting peak cladding temperatures. Both temperatures

, are acceptable and are within the AEC Interim Acceptance Criteria. These results show that there is no adverse effect of the insert for large breaks.

For small breaks, the core flooding tanks provide water at a very slow rate. Thus, the important parameter in the core flooding tank system is its pressure volume relationship and not line resistance. An increase of only 3% in CFT line resistance would have no effect on small breaks.

I III-2

a

~ FTCURE 1 CORE FLOODING TANK INJECTION RATE DURING AN 8.5 FT '

COLD LEG LOCA FOR VARIOUS FLOW RESISTANCES 4 8000 1

7000 f 4,

\.

Experimental l k = 5.5 _.--

+ insert 6000 rr  %

y/

3,,, t f

\

\ 's %g i = 8.3 _ _ _ _ Oesign

+ Orif. ice j  % .

y, 4000 f \ m w

\ w I

N s\ %\ ..

E 3000 k s 2000 1000 0 '

O 4 8 12 16 20 24 28 32 36 40 Time s

FIGURE 2 HOT SPOT CLADDING TEMPERATURE DURING AN 8.5 FT COLD LEG **

e LOCA FOR VARIOUS FLOW RESISTANCES IN THE CORE FLOOD TANK LINE 2600 2400 2200 a -~~~~-~~____

g

~*.

o' s e s~3 f [

( ~

N

'~

3300 f

• 1600 N

1 I 5

• 1I10 0 _ . _ _ K = 5,5 ( 2116*F)

?

4;

I200 G

1000

- - _ _ _ K = 8. 3 ( 216ttop) 800 600 0 10 20 30 40 50 60 70 T im e, s

_ _ _ _ _ _ _