ML20002C999

Test Design & Analysis Bulletin,Brp II,Re-Evaluation of Potential LOCAs at Big Rock Point

ML20002C999
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Site:	Big Rock Point File:Consumers Energy icon.png
Issue date:	11/10/1967
From:	Allred C, Bray A CONSUMERS ENERGY CO. (FORMERLY CONSUMERS POWER CO.)
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ML20002C998	List: ML20002C997 ML20002C999
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NUDOCS 8101150709
Download: ML20002C999 (41)

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ATTACIIMENT A TEST DESIGN AND ANALYSIS BULLETIN BRP II Re-evaluation of Potential Loss of Coolant Accidents At Big Rock Point November 10, 1967 Prepared by:

C. D. Allredl Tes.t Design and Analysis Unit Field Engineering Subsection Reviewed by:

A. P. Bray, Manager Systems Conformahce Unit Systems Engineering Subsection 7J4LXP7&f

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TABLE OF CONTENTS SPCTION PAGE I.

SUMMARY

1 II.

INTRODUCTION 2

III.

GENERAL CONSIDERATION 3

A.

Containment Integrity 3

B.

Fuel Rod Integrity 4

C.

Nature of a Postulated Repture 5

1.

dreak size 5

2.

Break Location 5

3.

Opening Rate 6

D.

Fuel Bundle Lift Forces 6

IV.

. ANALYSIS 8

A.

Blowdown Analysis 8

B.

Heat-Up Analysis 9

C.

Pipe Break Cases Analyzed 9

1.

Top Break Spectrum (Region I) 10 2.

Bottom Break Spectrum (Region II) 10 V.

RESULTS AND DISCUSSION 11 A.

Top Break Spectrum (Region I) 11 1.

Breaks Inside Enclosure 11 2.

Steamline Breaks outside Enclosure 12 B.

Bottom Break Spectrum (Region II) 12 References 14 i

Table I 15 Tr.Sle II 16 Figures I - VI i

I.

SUMMARY

The Consumers Power Company has engaged CF-APED to conduct a series of analyses considering potential loss of coolant accidents for a safeguards re-evaluation of Big Rock Point Reactor. This report contains the results of that study.

The analysis covered a spectrum of rupture areas at various locations in the system in an effort to establish the limiting break which the existing system could sustain without melting of fuel cladding.

Instantaneous, complete ruptures were assumed to occur, even though experience indicates that breaks would be at the worst partial cross section.

The system was divided into two regions, above and below the core, and antlyses of breaks were performed accounting for important effects peculiar to each region. The results of the study indicate that the existing Big Rock Point Reactor could tolerate, without fuel melting all sizes and locations of breaks which could reasonably be postulated if feedwater addition continued throughout the incident. However, if feedwater were lost and a rupture equivalent to a two-inch line or smaller were to occur, the slow pressure decay would prevent operation of the core spray until after melting begins. Further engineering, as well as economic analyses, should be made to determine the best measures which would decrease the probability that primary ruptures will result in cladding melting.

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- 11.

INTRODUCTION The rupture of any primary system piping of a nuclear power reactor could present a radiological hazard.

Should the cort olant be lost from the pressure vessel and not restored and main;ained, the heat generated by fission product decay could heat the fuel rods to the point of clad perforation and eventual melting.

This would release radioactive materials from their fundamental containment, the fuel rods, and could increas' the chances of contamination release to the atmosphere.

The Consumers Power Company has engaged GE-APED to re-evaluate the consequences of potentially hazardous loss-of-c. slant accidents which could reasonably be postulated at Big Rock Point. This report presents the results of the loss-of-coolant re-evaluations in response to recent AEC interest.

The analyses evaluate the response of the Big Rock Point plant to a spectrum of postulated ruptures in the primary system.

The results allow the assessment of potential hazards, and hence, the adequacy of the existing emergency core cooling provisions.

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, III. GENERAL CONSIDERATIONS A.

Containment Integrity Should the integrity of the fundamental containment barrier represented by the primary system and the fuel rod cladding be lost due to a primary system rupture and subsequent fuel melting, the enclosure sphere is designed to contain the radioactive fission products and the hot fluid contents of the primary system. However, the enclosure is not designed to withstand the ef fects of a gross core meltdown.

The containment design parameters and dimensions extracted from reference 1 are listed in Table I.

If a minor part of the fuel cladding were to melt as the result of a loss-of-coolant accident and fall to the bottom of the reactor pressure vessel, releasing gaseous and volatile fission products to the enclosure, the release of radic-active contaminants to the environs would be prevented.

However, if a major portion of the cledding were to melt, eventually the pressure vessel would be penetrated by the molten cladding fuel mixture which would be kept in the molten state by fuel decay energy.

If the molten core material were to penetrate the reactor pressure vessel, the eventual termination of the accident is not certain.

Since the continued integrity of the containment cannot be assured, the only way to assure that radioactive materials do not escape to the environs is to prevent clad melting.

A recent evaluation of the effects on the enclosure of a major accident are described in reference 2.

The accident defined, the so-called " Maximum Credible Accident," is the instantaneous, complete severance of the two twenty inch bottom inlet lines to the vessel while the reactor is in the hot standby condition.

In this condition, water with a higher energy content per unit volume would occupy space which would normally be occupied by vapor if the reactor were at power. The release of this greater amount of mass and energy to the enclosure sphere therefore results in a pressure transient which is more severe than if the reactor were at maximum rated power of 240 MWt.

The peak enclosure pressure resulting from the blowdown is calculated to be approximately 20 psig. This value is reached sixteen seconds after the postulated rupture when the reactor vessel pressure has decayed to a point at which core spray cooling is available.

It is noteworthy that the peak pressure is only 29.0 psig even if the rea assumedtooccurinstantaneously.tsfrcThis value therefore re-presents an upper bound for the peak pressure compared to the design value for the enclosure of 27 psig, and it is surely within the factor of safety inherent in the applicable ASME design : ode.

After its peak value is reached, the behavfor of the pressure is dictated by the net effects of heat losses and additions. Heat losses occur to the cooler structures within the enclosure, to the enclosure shell itself and to the environs.

Fission product decay heat generation, reactor vessel and piping sensible heat, and, if it occurs, metal-water-reaction result in heat addition to the enclosure. At the end of the b1cwdown the pressure reaches its peak value and then decreases as the cooler structures and steel shell absorb energy at a greater rate than it is added. After several minutes, however, the rate of heat-addition exceeds the rate of heat absorption as the temperatures of the cooler materials increase. Therefore, some means must be provided to prevent the pressure from exceed-ing its original peak value. This is accomplished by use of the post-incident system. A further purpose of this system is to accelerate the pressure decay, minimizing the leakage of any radioactive fission products, which might be present, through any imperfections in the enclosure shell to the environs. The post-incident system is supplied by the fire protection system pumps. Two independent branches from the fire protection system, one serving as a backup to the other, lead to two separate spray headers located in the upper part of the sphere. Operation of the post incident spray is calculated to reduce the containment pressure to slightly above atmospheric about six hours after the postulated "PCA".

B.

Fuel Rod Integrity An important objective of a loss-of-coolant analysis is to assess the amourt of damage to the reactor fuel which might result from a postulated rupture in the system. This being known, the attendant potential hazards can then be evaluated.

Should the reactor core become uncovered, the rate of heat transfer from the fuel-rods would be drastically reduced since the only modes of cooling would be convection to an atmosphere of steam, and radiation to cooler materials.

Except for this convection and radiation, the heat generated by fission product decay would remain in the fuel rods themselves and their temperatures would increase unless some external means of cooling is previded.

The Big Rock Point Reactor is equipped with a core spray cooling system to prevent melting of the fuel rod cladding should a loss-of-coolant accident occur.

Initially the system is supplied from the fire protection cystem, as in the case of the post-incident system, followed several hours after the rupture by recirculation from the bottom of the sphere through the core spray pumps and heat exchangers back to the reactor vessel. The core is sprayed from above by water from a circular sparger with nczzles which distribute the coolant among the fuel elements.

i C.

Nature of a Postulated Rupture The response of the system to a loss-of-coolant accident and the resultant severity of the events are highly dependent upon the nature of the break. This includes such characteris-tics as the size of the rupture, its location in the system and the rate at which the opening occurs. A schematic of the BRP primary system is presented in Figure I which shows the reactor vessel, steam drum, the risers, downcomers and other important piping and connections.

Important pipe sizes and the relative elevations of the vessel, core and dtum are given.

The primary system is also divided into two regions as an and for categorizing the problems to be analyzed.

1.

Break Size The size of the opening establishes, in part, the rate of mass-loss.

Since the rate of mass-loss decreases with system pressure, the makeup provided by feedwater could eventually exceed the rate of loss 12 the break is small enough, and reflood the core. The effect of break size on depressurization time is also very bnportanc because the core spray cannot provide cooling flow until the pressure has been reduced to a low value. Generally depressuri-zation is more rapid the larger the break because energy is removed faster. Hence core spray cooling can be provided sooner af ter the time that the core is un-

covered, i.e., when fuel heat-up begins, For smaller breaks, the time interval between the beginning of heat-up and depressurization might be sufficient to allow melting to begin before cooling can be provided.

2.

Break Location I

Break location is also an important parameter. Ruptures above the core are less severe than those below for two reasons: The system depressurizes faster and, once depressurized, the only loss of coolant and, hence, make-up requirement is due to a small amount of atmospheric boiling. Depressurization is retarded by flashing of hot liquid to vapot, but a top break vents off this vapor, while a bottom break does not, and removes energy at a high rate as well.

For example, if a bottom break were to occur, all of the liquid coolant could be eventually expelled, uncovering the core and leaving the primary system filled with vapor, still at an elevated pressure.

On the other hand, for a small top break, the blowdown could be entirely vapor, and the mass and energy removal would be such that the system would reach the depressurized state with a major portion of its liquid inventory re-maini Reference 5 shows that for the hypothetical case

of a vessel filled with saturated liquid water at 1000 psig, after a purely vapor blowdown to atmospheric pressure with no heat or mass addition, approximately 60% of the original liquid would remain. Calculations of the BRP 4

system show that a purely vapor blowdown through a six-inch line, e.g., the emergency condense. inlet line, would leave approximately 42%.of the original liquid volume, enough to keep the core cooled. O,the other hand, should a large instantaneous rupture occur high in the system, the rate of pressure decay would be so great that the extensive flashing to vapor of the ' system liquid would tend to fill the system with a homogene3us two phase mixture.

As a result, large amounts of liqri-d would flow out the rupture even if it were above the normal water level such as the severance of the twelve-inch primary steam line.

3.

Opening Rate Also of importance concerning the nature of a rupture is the rate of which the break opeus.

Since a break is merely a postulated event with rnany possible causes and mechanisms, to specify the timo for a pipe to rupture would be largely conjecture. Hence, ruptures are conservatively assumed to occur instantaneously, resulting in a greater loss of system mass for a given final break size and more liquid entrainment for a top break. General utility experience and that gained from the small cracks discovered recently at the Dresden BWR in small diameter seamless piping indicate that such crack propagation is quite slow and that a very small but detectable leak will appear long before any catastrophic failure. Hydrostatic testing, measurement of airborne radioactivity and ultrasonic inspection are among the methods which can be employed for early detection of such cracks. Therefore, one would more realistically expect not an instantaneous, complete failure of a given pipe but, at most, a slowly propagating opening, characteristic of ductile failures, the area of which would be only a portion of the total pipe cross sect. ion. Time would therefore be availabic icz early de-l tec. ion and for a decision to shutdown the reactor, if necessary, or take other appropriate measures.

D.

Fuel Bundle Lift Forces For top breaks, the pressure drop across the fuel bundle in-I creases as the pressure in the upper portion of the system decays and the upward flow accelerates.

This i

beenpreviouslyanalyzedfortheERPreactor.(gyoblemhas While the fuel bundles weigh in excess of 400 lbs., reference 6 reports that the maximum expected lif ting force on a fuel bundle would be approximately 200 lbs. The worst possible break, a l

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_7 complete instantaneous severance of one of the four risers was assumed to occur while the reactor was operating at 240 MWt.

Furthermore, the system pressure was assumed to be 1500 psia, yielding greater blowdown rates than would i

occur at the normal 1350 psia, and a fuel bundle was assumed to become completely clogged. Both of these assumptions result in larger than anticipated lift forces. Based on this analysis, lifting of the fuel would not be expected to occur for even the largest possible break.

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, IV.

ANALYSir Pipe ruptures of a catastrophic nature are not expected to occur for the reasons discussed earlier.

However, the consequences of a range of break sizes at various locations in the system were evaluated to determine the limiting rupture size which the existing system could tolerate without meltind of fuel cladding.

This portion of the problem has two facets:

a blowdown calculation and fuel heat-up calculation.

A.

Blowdown Analysis The theory of the analytical model used to describe the system is outlined in reference 4.

The equations presented are solved by means of a digital computer assuming that the system responds as a single pressure node.

Basically the model calculates the loss of mass from the break and the pressure transient. It accounts for flashing due to pressure decrease and loss of fluid from the break utilizing blowdown rates from reference 5 and determines the remaining liquid inventory.

From the amount of liquid remaining, one can then determine whether the reactor core has been uncovered and at what time.

If the core does become uncovered, this information can be used to obtain the time which exists to provide cooling before the commencement of meltdown.

If the core is cooled before the fuel cladding begins to melt, the existing system can tolerate the postulated rupture.

The reactor is assumed to be at rated power of 240 MWt, the maximum value allowed by the license, and an instantaneous rupture of a given area is assumed to occur.

The model predicts the behavior of the system simulating the effects of the following phenomena; (1) single or two phase blowdown of the system; (2) the addition of feedwater; (3) cool ~

down of the rhedweter heaters; (4) the shutoff of steam flow by the turbine admission valve under control of the pressure regulator; (5) sensible fuel rod heat and 1: eat generation due to delayed neutrone and fission product decay; and (6) core spray cooling as a function of reactor vessel pressure.

A given accident is assumed to run its course approximately an follows:

(1) Blowdown of the system commences immediately.

(2) The reactor is shutdown very quickly by increased moderator void volume and by rapid control rod insertion due to low drum level and high containment sphere pressure.

(3) As the pressure decays the pressure regulator closes the primary admission valves to the turbine, isolating the reactor from the main condenser.

In most cases, this would occur faster t'an the automatic closure of the steam line sphere isolation c

valve which requires about sixty seconds after actuation.

(4) When the pressure reaches approximately 140 psig, core spray flow begins and increases as the pressure decays further. Whether cooling is provided before any melting of

_9 the fuel cladding occurs depends upon the break size, its location and the addition of feedwater.

B.

Heat-Up Analysis The model employed to analyze the heat-up phase of the transi-ent is described in reference 4.

For the calculation the reactor core was divided into radial and axial regions.

The fuel bundles were further subdivided into zones of fuel rods to allow for the effects of rod position within a bundle.

The model accounts for radiation heat transfer between fuel rods and to the cooler channel walls. The amount of hydrogen released by the exothermic metal-water reaction and the thermal energy carried away by the hydrogen are also calcu-lated. This calculation assumes that an unlimited amount of water is available to react, a very conservative approach. The remaining heat, whether generated by metal-water reaction or by fission product decay, is assumed to remain in the fuel rod. The assumptions employed for the commencement of heat-up are consistent with reference 4.

The core does not have to be covered with liquid to be effectively cooled because of the agitation associated with the boiling process. Based on conclusions from reference 4, this analysis assumed that for most breaks heat-up began when the core was one-half uncovered. However, if the rupture is quite large, e.g.,

the severance of a recirculation line, experiments at APED (4) indicate that high heat transfer coefficients in the nucleate boiling range exist for only a short time after the break while the liquid film on the fuel rod evaporates.

Then the heat transfer deteriorates rapidly as the fuel be-comes steam blanketed, and heat-up begins even though the core region is still flooded. Heat transfer coefficients inversely proportioned to time after dryout were established, and these values were employed in this analysis for the recirculation-line break.

C.

Pipe Break Cases Analyzed Since dif ferent phenomena predominate, distinct analyses were performed for a range of break sizes depending upon the location of the postulated rupture.

The primary system was divided into two regions as shown in Figure I.

Region I is the portion of the system above the reactor core in which breaks are characterized by vapor or liquid-vapor blowdown with depressurization generally occurring at less expense to the liquid inventory than for bottom breaks. Region II contains the portion of the system below the core in which the blowdown is either liquid or a liquid-vapor mixture.

, Because there are several sizes of pipes which could be pcstu-

. lated to rupture either partially or completely, a range of break creas was investigated in each region to determir.e the limiting size which could be rolerated without melting of the fuel cladding.

1.

Top Break Spectrum (Region I)

Table II lists the pipes in this region and their sizes.

In addition to breaks within the enclosure, the possi-bility exists for ruptures in the steam line after it has penetrated the containment shell.

Primary fluid would blow into 'he turbine building and subsequently reach the atmosphere until the steam line isolation valve is closed. A spectrum of postulated breaks in the steam line outside the enclosure was analyzed which included the complete severance of the steam line.

The most severe ruptures which can be postulated above the core are those of large area which result in two phase flow out the break, i.e.,

greater liquid loss than a pure vapor blowdown.

Heat-up was assumed to commer e when the amount of liquid remaining in the system was insufficient to cover the reactor core to its mid-plane. This is a conservative approach since cooling of the core would in reality be provided for a longer time by the two phase mixture which would still c'over the core.

The analysis was performed in the above manner for postulated ruptures inside the enclosure and those of the steamline outside the 1 closure.

However, for steamline breaks outside ti.e enclosure, the isolation valve was assumed to be actuated thirty seconds after the rupture and to be fully closed sivty seconds later, isolating the break from the reactor. The time delay before actuatio* is for the reactor operator to manually initiate closure of the valve after observing an abrupt decrease in reactor pressure and coincident alarms from the turbine building area radiation monitors, i.e.,

no credit is taken for automatic actuation which would occur earlier due to low reactor water level.

2.

Bottom Break Spectrum (Region II)

Table II lists the pipes and their sizes which could be postulated to rupture in Region II.

Because ruptures in this location are below the ccre all sizes of breaks will eventually drain the system if no make-up is pro-vided. Therefore, the ability of the existing system to tolerate any such rupture depends upon the core spray system to be available for a considerable period of time af ter the incident.

. V.

RESULTS AND DISCUSSION A.

Top Break Spectrum (Region I) 1.

Breaks Inside Enclosure The analysis of possible breaks in this region indicates that the existing Big Rock Point Reactor could sustain, without melting any fuel cladding, a rupture as large as that equivalent to the complete, instantaneous sev-erance of a seventeen-inch downcomer, the largest credi-ble break for this case. This is illustrated by Figure II.

Curves were drawn for the entire spectrum to show the time at which the midplane of the reactor core would become uncovered, the time at which adequate core spray flow would be possible, and the time at which melting would commence if cooling were not provided.

The assumed break area relates to the complete severance of two equivalent pipe sizes, dependent upon whether the postulated rupture is single or double-ended, i.e.,

if primary fluid reaches the break via both parts of the severed pipe.

For example, the complete severance of the shutdown suction line would be a single-ended break while that of a fourteen inch riser would be double-ended.

A core spray flow of 250 gpm, which would occur at a reactor pressure of 85 psia, was assumed to provide ade-quate cooling. This assumption is based on current designpracticeanddistributionmeasurementswhigg)were performed with a mockup of the BRP corfiguration Figure II also shows the effect of rated and zero feed-water addition during the accident. While the times to uncover the core midplane, i.e.,

initiation of heatup, are not significantly affected because the blowdown is so fast, feedwater addition causes the reactor pressure to decay faster and makes the core spray available sooner.

Since it is possible that the feedwater pumps could over speed and trip off as the pressure decays and the level regulator, seeing a falling level, signals for more flow by fully opening the control valve, the case for zero as well as rated feedwater addition after the break was analyzed. For this region, however, loss of feedwater has orJy an insignificant effect on those breaks for which the core becomes uncovered.

Heatup calculations were performed for the fuel currently in the BRP reactor which has an 11 x 11 array and also for the Type C reload fuel (9 x 9 array) which will be added later. The 9 x 9 fuel is shown on Figure II to begin melting slightly earlier than the current fuel if no cooling is provided. This is primarily because the Type E fuel has a greater diameter; consequently there is more stored heat and power generation per fuel rod. Of further interest are the maximum cladding temperatures and percent of perforated rods with core spray cooling.

Figure IV shows these parameters for both types of fuel, and the 9 x 9 fuel is found to reach greater temperatu. e', than the 31 x 11 fuel. However, for break areas less than 0.8 ft, fewer perforated rods occur in the Type E fuel because of the number and locations of peaked rods.

2.

Steamline Breaks Outside Enclosure The analysis of possible breaks in this region indicetes the existing system could sustain, without melting any fuel cladding, a rupture as large as that equivalent to the complete, instantaneous severance of the twelve-!nch steamline, the largest possible break for this case.

The same break area versus time curves as discussed for breaks inside the enclosure are presented in Figure III.

The closure of the steam 1(ne isolation valve, assumed to occur 90 seconds after the break, is found to retard the pressure decay such that full operation of the core spray is slightly delayed. However, the time to uncover the core is unaffected by the valve closure since this would always occur before the valve was assumed to be actuated. The maximum temperatures and percent of perforated rods are also shown on Figure IV.

B.

Bottom Break Spectrum (Region II)

The analysis of postulated breaks in this region indicates that the existing system could sustain, without melting any fuel cladding, a rupture as large as that equivalent to the complete, instantaneous severance of a twenty-inch recirculation line, the largest credible break for this case.

However, in order to sustain all breaks, feedwater eddition is necessary. As seen on Figure V fuel clad melting begins before the pressure is low enough to allow adequate core spray ficw for small breaks since the time-to-melt curve crosses over the core spray curve. This can also be seen on Figure VI where the cladding temperatures reach melting 2

for both types of fuel for breaks between 0.02 and 0.03 f t,

This would correspond to the complete severance of pipes slightly greater than two inches in diameter or, of course, ptrtial breaks of larger pipes.

From Table II, it is L

I seen that all major piping in this region could be postulated to break in such a manner to cause fuel melting.

Feedwater addition serves to delay uncovering of the core, delaying heatup, and to accelerate the pressure decay, allowing sor ar core spray operation.

If feedwater cannot be assumed.ther measures would bc necessary.

Further engineering and economic analysis should be performed to determine the optimum methods.

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REFERENCES:

1.

Appendix "A", Consumer Power Company Big Rock Point Nuclear Plant Technical Specifications, Appended to Operating License Number DP-6, reissued November 15, 1966.

2.

Change 8 to the Big Rock Point Nuclear Plant Technical Specifications, December, 1965.

3.

F.J. Hoody, " Enclosure Pressure and Temperature Corresponding to Additions and/or Extractions of Water Phases and Heat,"

GEAP 3315, August 22, 1960.

4.

Dyster Creek Unit ;4o.1 Facility Description and Safety Analysis Report, Volume II,' Appendices E, J, and K.

5.

F.J. Moody, " Maximum Two Phase Vessel Blowdown from Pipes," ASME Paper No. 65-WA/HT-1.

6.

F.J. Moody, " Fuel Lif t Forces Following Primary Steam Rupture,"

APED 3883, February,1962.

7.

E. Janssen, Letter to L.K. Holland, November 2, 1965.

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TABLE I CONTAINMENT SPHERE DESIGN PARAMETERS AND DIMENSIONS Design pressure, internal, psia 41.7 Design temperature rise, 'F 190 (Coincident with design internal pressure)

Design maximum temperature 235 Design pressure at minimum temperature, 41.7 maximum internal, psia Wind load ASA Std. A58.1 Without snow load (Basic wind pressure =

30 psi)

With snow load 60 mph Snow Load ASA Std A58.1 (Max. = 40 psf at top)

Laterial seismic acceleration, percent of gravity (coincident with dead load and snow 5

load only)

Design external pressure, psig 0.5 (not limiting, safe external pressure is 1.22 psig)

Design maximum ambient temperature, *F 130 Design minimum ambient temperature, 'F 45 Permissible air leakage rate 41.7 psia at ambient temper ature, percent per day of 0.5 free volume (including all penetrations)

Diameter, ft.

130 Height above grade, ft.

103 5

Approximate f ree' volume, f t 9.4 x 10 i

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. TABLE II BRP Primary System Piping Break Size Break 2

Region Description No.

Type in.

Area,ft Single ( )

1-1/2 0.0126 1 (above Vessel to drum vent 1

• • "E

1. • }

Poison Tank pressure 1

Single 2

0.0205 Core Spray 1

Single 3

0.0458 Single (3)

Emergency Condenser 2

4 0.0798.

Condensate return Single ( )

Emergency condenser inlet 2

6 0.131 Primary steam 4

Double 8

0.634 branches 2

Double 12 1.268 Feedwater branches

.2 Double 8

0.634 Shutdown Suction 1

Single 8

0.317 Feedwater 1

Single 10 0.498 Steam 1

Single 12 0.715 Riser 6

Double 14 1.704 Downcomer 4

Double 17 2.546 II (Below Reactor Core)

Poison 1

Single 3

0.0458 Cleanup Suction 1

Single 3

0.0458 Cleanup return 2

Double 3

0.0916 Thermal equalization 1

Double 4

0.1596 Crosstie Single ( }

5 0.1264 Recirculation Discharge 2

Bypass Shutdewn return 1

Single 6

0.181 Downcomer 4

Double 17 2.546 Recirculation 2

Double 20 3.506 (1)

Maximum possible value assuming complete severance (2)

Double-ended if valve open (3)

Double-ended if emergency condenser in operation f

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,s.. *..,, f ATTACHMENT B BIG ROCK ~ POINT REDUNDANT CORE SPRAY SYSTEM t 1.0 General The-original' Big Rock Point post incident cooling and core spray cooling systems are described in the Final Hazards Summary Report, Sections 3.7 and 5.9 respec:ively, and as shown on drawing M-123, Rev. 8. The original systems each consisted of essentially single systeme with initial water supply from the plant fire protection system followed by water supply-from the core spray recirculation system. ~ Each water supply system has two full capacity pumps which are arranged to supply water ttrough an automatic operated single valve set to the core spray or through an automatic single. valve to the post-incident spray nozzles. The back-up enclosure spray is provided with a.normally closed manual operated valve. .New criteria requiring that safety related systems be designed with redundancy to meet the single failure criteria for active components has resulted in required modifications of the two system 9. ?!he modifications e.te covered below 9.feEXMM ) and as shown on (.rawing M-123, Rev. ~2.0 Core Spray System 2.1 A second core spray system has been added and is in parallel with the existing spray system. The new core -

~ -~ i. spray system is redundant to the existing core spray system. The water supply to the redundant spray system can be remote manually selected from either of the two existing-fire system supply headers or from the core spray recirculation system. Selection of the supply source can be accomplished by operating a new a-c motor operated valve, MO7069, cross-con-necting the two fire system supply headers. The new spray system valvas are a-c motor operated and are therefore redundalt to the existing d-c operated valves. Operatio. of the new spray system valves is automatically initiated by primary system level and pressure switches which have been added and are separ-ate and redundant to the control switches for the existing spray valves. 2.2 Another source of water from the fire protection system can be remote manually routed from the core spray heat exchanger shell inlet to the channel outlet which supplies either of the two core spray systems. Heat exchanger bypassing is accomplished through a d-c motor operated valve MO7072. 2. All piping and valves added for the redundant core spray system are manufactured in accordance with appropriate sections of USAS 31.7 and the ASME Draft Code for Pumps and Valves for Nuclear Service. '

'I J l 2.4 All electrical wiring for actuation of the redundant core spray system is routed in metal conduit and thereby kept physically and electrically separate from the existing spray system electrical system. The power supplies-for the new valves are shown on Single Line Diagrams E-102, Rev. 4, and E-101, Rev. 8. 2.5 All piping sections in the enclosure spray system, spray recirc system and core spray system will be missile protected where required to prevent a single failure from causing total loss of function of the affected system. Missiles considered are those or-iginating from the high pressure reactor system. (Spray nozzle details by GE) 3.0 Post Incident Cooling 3.1 The existing post incident cooling system consists of one automatic actuated spray and a back-up enclosure spray which is actuated by manually opening of a nor-mally closed gate valve outside of the containment. The post incident cooling system has been modified by the addition of an automatic a-c motor operated valve in the back-up spray header. This new valve is automatically opened by one of two containment high pressure switches. Water supply to either of the two full capacity spray systems can now be aligned from I P either of the two fire system headers or from the spray recirculation system. 3.2 The existing enclosure spray valve has been modified from a-c motor to d-c motor operated. This arrange-ment now insures that one enclosure spray system and one core spray system will remain operative in the event of either a-c or d-c power failure and the operative systems can be supplied from either the fire system or the spray recirculation system. 4.0 Miscellaneous Modification 4.1 Interconnection of the existing systems has required that containment isolation check valves be added in the spray recirculation return lino-and the fire header which initially supplied only the back-up enclosure sprays. 4.2 A flow indicator and switch for annunciation on low-flow has been added to the existing core spray system and to each enclosure spray as well as the new redundant core spray system. The indicators and alarms are located in the main control room. 4.3 Two duplex type basket strainers have been added to l the fire protection system. These strainers are lo-cated to insure that all fire system water into the containment sprays and reactor sprays does not contain..

solids that could plug a-spray equipment. Each strainer is equipped with a local differential pr' -ure indicator and high differential and no-flow a: lo-cated in the main control room. 5.0 Summary of Modifications 1. Addition of a 4 inch core spray header originating from the existing back-up enclosure spray line within the containment. This line contains two 4 inch a-c Motor operated valves (MO-7070 and 7071) and one 4 inch check valve similar to the existing core spray valving. 2. Addition of a 4 inch a-c Motor operated valve (MO-7069) cross-connecting the existing core spray line to the new core spray line. 3. Addition of a 4 inch a-c Motor operated valve (MO-70 6 8 ) in the existing back-up enclosure spray line. 4. Addition of a 4 inch d-c Motor operated valve (MO-7072) cross-connecting the fire system supply to the shell side of core spray heat exchanger to the spray recircu-lation line. 5. Addition of two 4 inch containment isolation check valves in the existing back-up enclosure spray line and the spray recirculation l'.ae. c5-

6. Addition.of a bypass with a 4 inch manual valve around the existing valve, MO-7066, in the fire ~ system cooling water supply line in the core spray heat exchanger. 7. Addition of two duplex strainers (BS-5761 and 5760) in the fire protection system; one in the alternate flow path off the_outside fire loop, the other in the fire system supply line to the core spray heat ex-changer. 8. Addition of four flow indicators, and low-flow alarms, one in each of the two core spray lines and one on each of the enclosure spray lines. 9. Addition of two reactor pressure and level switches to control the new core spray valves. 10. Addition of two containment high pressure switches to control the new back-up enclosure spray valve. Failure Analysis The modified core spray and post incident cooling systems faulure analysis is summarized on Table 1. i l

^ TABLE I BIG ROCK POINT PLANT CORE SPRAY SYSTEM - FAILURE ANALYSIS COMPONENT MALFUNCTION COMMENTS AND CONSEQUENCES 1. Diesel Fire Pump Fails to start Electric Driven Fire Pump from Emergency Diesel Generator will deliver required flow. 2. Electric Fire Pump Fails to start Diesel driven fire pump will deliver required flow.~ 3. Strainer BS 5759 Pluqqed Alternate flow path to' spray system from yard outside loop will provide required flow. 4. Core Spray Valve MO-7051 or 7061 Fails to open or d-c Required flow i' available through the redundant (dc-Motor operated) power failure core spray syst 4 5. Core Spray Valve MO-7070 Fails to open or Recuired flow is avillable through the redundant or 7071 (a-c Motor a-c power failure core spray system. operated) 6. Containment Pene Sticks closed Required flow is available through the second tration check fire supply line supplemented by the cross-connection valve in either fire from the recirculation piping. system supply header. 7. Core spray line check Sticks closed Required flow is available through'the second valve in either spray core spray line. system. 8. Core spray Ring or Plugged Either core spray system provides 100 percent of core spray nozzle. required flow. 9. Recirculation Suction Plugged Sufficient capacity is available to ensure adequate Strainers. flow in the event that 3 of the 5 strainers become plugged. 10. Core Spray Heat Exchang Fails to open or 1. Manually operated bypass valve will allow full er Cooling Water Supply are power failure rated flow through the core spray H.X. valve MO-7066 (a-c 2. Bypass around the H.X. (MO-7072) will allow required flow to enclosure soray or core spray motor operated) ring

CORE SPRAY SYSTEM - FAILURE ANALYSIS (Continued) COMPONENT MALFUNCTION COMMENTS AND CONSEQUENCES 11. Cross-Connecting valve Fails to open or a-c 1. Required flow is available through either between fire system power failure. core spray system and either containment supply headers MO-7069 spray header from the ff.re system. (a-c motor operated) 2. Required flow is available through the enclosure spray headers or the core spray ring from the spray recirculation system. 12. Valve cross-connect-Fails to open or ing the Fire System to d-c power failure Required flow is available through the new the Recirculation core spray or back-up enclosure spray. Piping MO-7072 (d-c motor operated) 13. Containment Isolation Sticks closed Required flow is available through the existing Check Valve in Recir - core spray or containment spray headers. The culation Line spray recirc system is provided with a test tank and piping arrangement to permit periodic system testing. Recirculation through the' check valve during testing will insure valve operability. 14. Reactor level-pressure Electrical failure Required capacity flow is available through the control circuit to second spray system controlled oy redundant either core spray level-pressure switches. system. 15. Containment pressure Electrical Failure Required capacity flow is.available through the control circuit to second spray system controlled by redundant pres-either enclosure sure switches. i spray system generator through %iier#gle from emergencyteY."* Requ D-C power supply pvaila 16. A-C power supply Failure available through the enclosure spray and through the core spray ring. 17. D-C power supply Failure A-C power supply available from station power transformer 11. Requi ed flow is available through the redundant core spray nozzle and through the back-up enclosure sprays.

O ATTACHMENT C December 4, 1969 MEM0RANDUM REDUNDANT CORE SPRAY SYSTEM ^ for BIG ROCK POINT PLANT CONSUMERS POWER COMPANY I. CONCEPTUAL DESIGN PARAMETERS, A. Conceptual Design Layout The attached drawing (731E748 sketch) defines the present concept of the envelope, hardware configuration, and piping interface for the core spray nozzle assembly. The essembly will consist of a single spray nozzle located centrally over the core and approximately at the elevation of steam bafile. Entry into the vessel will be thru the existing 10" noz-zle f.n the center of the vessel head. Several alternate concepts are presented for attachment of the core spray nozz7e assembly to the vessel head flange. Each concept shows a modified weld neck flange for the bolted attachment to the existing vessel nozzle. The interface with piping to be supplied by "cchtel is either a bolted connection to a 4" flange, or a weld joint to the 10" flange. In each case a thermal sleeve is provided for protection of the vessel flange joint from excessive thermal chock. Piping inside the vessel is designed for maximum stiffness for vibration considerations. The final selection of design details of the flanged joint will be deter-mined based on further analysis and evaluation of optimum joint geometry. The spray nozzle will be one of several possible nozzles manufactured by Spraying Systems Company or Spray Engineering Company. Some of the noz-zles being considered are Spraying Systems Catalog Nos. 2-1/2 R9590, 3H120, and Spray Engineering Company Catalog Nos. 13D or 13DN. Final de-cision on the nozzle to be used is pending receipt of further definition of core spray system flow and pressure parameters; and receipt of nozzle performance characteristics from the nozzle manufacturer. The attached Spraying System Company drawings are typical of the nozzle test data that will be used in selection of the nozzle and evaluation of nozzle performance, area coverage and distribution. System parameters used to arrive at a preliminary selection of nozzles were approximately 250 GPM @ 30 psi differential pressure and anproxim-ately 400 GPM @ 100 psi differential presaure. B. Design Analysis Vibration analysis will be performed on the assembly. The analysis will compare the lowest natural frequency of the spray pipe to the maximum

Memorr.ndum - Pegn 2-vortex shedding frequency associated with the highest possible flow around the pipe..The lowest natural frecuency at the pipe shall not be less than 1-1/2 times the highest vortex shedding frequency. Stresses due to seismic loading will be evaluated in accordance.with .GEISAR Appendix C. A 6" vertical displacement (jump) will be considered. Further definition of seismic loads and rate of vertical displacement for the 6" jump is required. The hardware and vessel attachment joint st.all be designed in accordance with the following codes and standards: 1. New pressure parts in accordance with-USAS B31.l'with analysis per - Taylor Forge Bulletin 502 (ASME Section VIII, Appendix II). 2. Evaluation of affects of bolted flange on vessel head nozzle per ASME Section III, N-415.1 and N-450. No detailed cyclic analysis will be done. 3. Flange bolting per ASME Section III, Appendix I-12. C. Perf ormance Testing Noz.tle selection will be based on performance data submitted by the noz-zie manufacturer. No performance testing is anticipated at this time. ,4 c$f/>ti s'd a J. A. Hallam' Reactor Assembly Engineering M/C 743 Extension 2042 ~ /cee i

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