Text

Responds to FOIA Request for GE Repts NEDE-10182 & GEAP-3143.Forwards Rept GEAP-3143, Test Rept for Pressure Suppression,Development Program
	ML20214K993
Person / Time
Site:	Humboldt Bay
Issue date:	05/26/1987
From:	Grimsley D; NRC OFFICE OF ADMINISTRATION & RESOURCES MANAGEMENT (ARM)
To:	Menz J; SIMPSON, THATCHER & BARTLETT
References
	FOIA-87-40 NUDOCS 8705290138
	Download: ML20214K993 (1)
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{{#Wiki_filter:f "'% UNITED STATES J 'k 0' NUCLEAR REGULATORY COMMISSION WASHINGTON, D. C. 20555 John R. Menz, Esquire Simpson, Thacher & Bartlett MAY 2 6 gg7 One Battery Park Plaza IN RESP 0 HSE REFER New York, NY 10004 TO F01A-87-40

Dear Mr. Menz:

This is in reply to your letter dated April 6, 1987 concerning Peter Thomas' Freedom of Information Act (F0IA) request designated as F0!A-87-40. With respect to the General Electric Company's (GE) report NEDE-10182, please be advised that Mr. Harold Neems of GE infonned my staff in February that the report could be provided to your law firm in response to the F0IA request but that the report could not be made publicly available because it contains information that GE considers proprietary information. With respect to GE report GEAP-3143, enclosed is a copy of that report which is located in the Humboldt Bay docket file. My staff contacted Mr. Neems, and he has informed us that this report does not contain information which GE considers proprietary information. Therefore, a copy of this report is also being placed in the NRC Public Document Room. The reproduction charge for this document is $7.45. With respect to the accession list for the pre-docketed material for Humboldt Bay, please be advised that the list provided in my letter dated March 20, 1987, was in fact prepared in response to the F0IA request. Sincerely, A,Mm t. Y '? Donnie H. Grimsley, Director Division of Rules and Records Office of Administration and Resources Management

Enclosure:

As stated g52 8 870526 ~ NENZ87-40 PDR

h ((1t9~ GEAP 3143 R 39 APE-15 ~ l.~..'. I" : 3 c.; f Y T/ _._/ w fu >s 21 'i TEST REPORT R)R THE PRESSURE SUPPRESSICH -r-DEVELOPMENT PROGRAM Prepared For PACIFIC GAS AND ELECTRIC COMPANY By W. L. Fiock l E. Janssen A. G. Steamer i I April 2, 1959 i g ?EEN TD ).E0h_ATO Y CE" R00%' 06 ggjiORY 02E FECD?Y GENERAL ELECTRIC '( ATOMIC POWER EQUIPMENT DEPARTMENT g&(. Tf F ~ i 1 ~ , d.,.g. SAN JOSE, CALIFORNIA un D" f x, ~. Q

/ rca u:c cr e* supLevaca cuev GEN ER AL $ ELECTRIC i ATOMIC POWER EQUIPMENT DEPARTMENT TECHNICAL INFORMATION SERIES TITLE PAer AUTHoes sueJECT CLAselFICAT1oM NO.GEAP 3143 W. L. Fiock Reactor Containment R 5Q APE-1 % E. Janssen OAT A. G. Steamer April 3, 1959 TEST REPORT K)R THE PRESSURE SUPPRESSICN DEVELOPMENT PROGRAM ^ " ' " ^ *

Pacific Gas and Electric Company sponsored a development program for the pressure suppression system of reactor containment. This program included a condensing test facility to investigate large scale steam injection equipment performance and a small scale transient test facility to obtain pressure response data for a complete system. The results of tests conducted with (cont'd below) o.t. class.

nernooucleLE COPY FILED AT No.PAGE. APED Library None The test results were highly successful with conclusioNa respect to verifying and extending the technology of the pressure suppression containment system. The condens-ing tests indicated that the steam injection equipment may be simple and that condensation is effective for a wide range of design conditions. The transient test facility provided data that confirmed the results of the analytical model and assurance that the integrated system functioned properly. Tests with noble gas tracers offer encouraging evidence that the pool is effective for retaining fission products. Abst ra ct-c ont 'd these facilities are presented in this report. By cutting out this rectangle and folding on the center line, the above information can be fitfod into a standard card file. For list of contente-drawings, photos, etc. and for distribution see next page (FM410 2). Pacific Gas and Electric Co. inronwATian encPAnED Fan A. M. Kennedy, R. Edin and operating personnel of P. G. & E. G. B. Bethards, W. L. hiock, E. Janssen, A. G. Steamer, of General Electric. l cou=Tensieuro Engineerino .ccTion (300) 129 San Jose, California eUILDING AND Room Mo. t o,,,,,, arm *lt S ( t *S7 3 1

5 - N s-s t ~ t m lHEEl Etat I. INTRODUCTION I-1 II. CONCLUSIONS 11-1 III. CONDENSING TESTS Purpose of Tests III-1 Test Program III-1 Description of Test Facility III-3 Test Procedure III-5 Discussion of Results III-5 Conclusions from Test Results IV. TRANSIENT TESTS Purpose of Tests IV-1 Test Program IV-1 Description of Test Facility IV-2 l Test Procedure IV-6 l Discussion of Results IV-9 l Conclusions from Test Results IV-15 V. APPENDIX A. Condensing Test Sample Data and Detailed Program l B. Methods of Analysis C. Transient Test Data D. Methods of Analysis E. Reference Drawings l

ERRATA SHEET Pace No. II-1 Line 1, omit "trans ". III-2 Line 7, change "were" to --was-. III-7 Line 7, after the period insert --Figure VA-10 is a temperature chart for the bottom horizontal injector run in which steam was released into the vapor space. (The temperature scale is labeled as megawatts.)-. III-7 End of fifth paragraph, after the period insert --The temperature chart for the 4 inch multiple injector is shown on Figure VA-ll. This chart is typical for the vertical injectors.-. III-8 Line 1, change " smooth" to --smoothly-. III-8 Fourth paragraph, last line, after the period insert --Figure VA-12 is a typical temperature chart obtained during the runs for investigating pool vibration.-. III-9 Line 7, change " wide" to --wise-. III-9 Second paragraph, last line, af ter period insert --The intensity of the vibrations increased with increased steam flow rates under all conditions.-. III-11 After paragraph numbered 4, insert --Refer to Figure VA-13 for a typical temperature record obtained with the compartment tests. It should be noted that only one thermocouple was used to record temperatures of the water in the compartment. It was connected to several recording points.-. IV Line 3, insert --system-- after " suppression". IV-4 Third paragraph, line 6, change "shart" to --sharp-. IV-5 Fourth paragraph, line 1, change 9p9rtion" to --portion-. IV-6 Fifth and sixth paragraphs, lines 4 and 6, respectively, change "tupture" to --rupture-. IV-13 Second paragraph, line 3, should read --of Figures VC-7 and VC-11 with Figures VC-15 and VC-9,-. IV-16 Third paragraph, line 10, change " air" to --aid-. Fioures V-A-1 to V-A-8 Line 4, respectively substitute --2 MILLISECONDS / DIVISION-- for "10 MILLISECONDS INCH".

I. INTRODUCTION The test results obtained from the pressure suppression development program provide conclusive evidence that-the concept is practical for reactor containment. A pressure suppression containment system in its simplest form is shown in Figure I-1. The reactor vessel volume, dry well volume and containment volume are labeled as volume (1), (2), and (3), respectively. The water pool for condensing the released steam is part of the containment volume (3). In operation, steam would be released from a rupture in the reactor pressure vessel or the priraary coolant system and flow into the dry well volume; the steam is discharged from the dry well, through vent pipes, into the water pool where it is condensed. The salient features of the system are low containment pressures and the entrainment of released fission products in the water pool. As with most new concepts there are many design considerations that can be resolved only with model or prototype testing. Two major test facilities were constructed to obtain the experimental informa-tion considered necessary for the evaluation of a pressure suppression system. Pacific Gas and Electric Co. sponsored a development program that included a large condensing test facility at the Moss Landing Power plant of Pacific Gas and Electric Co., and a scale _podel, transient fest facility at the San Jose Plant of General Electric Co. This report presents the results of tests performed at these facilities. The condensing test facility consisted of a large tank of water with steam headers that would facilitate discharge of steam into the water pool with full size vent pipes of various arrangements. The steam supply system permitted flow rates up to 100,000 pounds per hour. These tests provided data for design of the steam injection equipment i and for the arrangement of the water pool. The scale model transient test facility was built to represent a 50 Mw prototype power plant reactor enclosure with a volumetric l scale factor of 1000 to 1. The facility consisted of a pressure i vessel, dry well and a containment tank with a water pool. The 1 facility was designed with the flexibility necessary for obtaining l pressure response data with variations c ' the many design parameters. The results of this test were primarily used to confirm the mathe-( matical model for predicting pressure response of the dry well volume and to evaluate effectiveness of fission product entrainment. l l l l I l I-l l

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II. CCNCLUSIONS The results of tests conducted with the condensing test facility and trans-transient test facility were highly successful with respect to verifying and extending the technology of pressure suppression systems for reactor containment. Tests performed with the condensing facility were significant in that, of the hundreds of various tests, steam release could be detected in only three instances. None of these tests w:uld be considered as reasonable for design condition. A simple pipe of suitable diameter may be used as a steam injector to accomplish rapid and effective condensation. The depth of submergence of the discharge end is not critical for complete condensation with vertical injectors. With horizontal injectors it is necessary to have suf ficient depth of submergence (a few feet) to prevent steam release. Under certain test conditions, there were severe pressure fluctuations in the pool. A series of tests were run to determine the magnitude and frequency of these pressure fluctuations. The conclusions from these 0 tests were that for. pool temperatures less than 120 to 130 F the fluctuations were very small and that under any conditions the magnitude could be reduced by supports fastened to the injector. Tests run with an internal compartment to reduce the effective volume of water were important to determine injector-pool geometry relations. The conclusions from these tests were as follows : 1. The two dimensional representation of the pool by using the compartments was valid for determining pool geometry. 2. Some of the injector-pool gent.etry combinations were better than others from the consideration of pool surface deflections and also pool mixing. 3. The air injected into the pool at the onset of the event will not interfere with condensation. Further, the air that is slowly purged out of the dry well will not materially affect complete condensation of the steam. The results of the transient test facility have provided conclusive evidence that the pressure suppression system is effective in containing the energy release from a boiling water reactor incident. Furthermore, the pressure behavior of such a system may be calculated with a reasonable degree of accuracy, and the values are conservative when compared to the test values. The test results also indicate that such a system will probably be very ef fective for retaining any fission products that may be released with an accident. II-1

The most important single design consideration is the peak pressure in the dry well. This pressure determines the design conditions for the dry well vessel. In general, the test results agreed very well with analytical model results. There were, of course, deviations between the two due to various causes such as flow conditions at the pressure vessel discharge (orifice), and subcooled conditions in the pressure vessel. In all cases, the analytical model results had higher values for peak pressure and shorter time intervals to reach peak pressure than corresponding test conditions. The peak pressure in the dry well is determined primarily by the b,reak area (pressure vessel discharge conditions), the yent depth of submerger.ce and the dry well volume. The depth of submergence is the easiest to control by design and various schemes may be used to keep water out of these vents. With no water inside of the vents, the peak pressure is reduced substantially. The vent area (above a certain minimum value for given conditions) has i little or no' influence on the peak pressure of the dry well. The vent area does affect the after-peak response of pressure in the dry well. It appears that this area may be reduced from that originally contemplated. Dry well design pressures were noted to have high values of negative pressure at the end of the event. It will be necessary to consider this behavior when designing the dry well. Another design consideration is that of pressure in the containment volume. The air expelled from the dry well enters the containment volume and compresses the gas in the containment volume in what appears to be an adiabatic process. The larger the ratio of containment volume to dry well volume the lower this pressure will be. l A significant feature of the entire system is that all of the respective volumes return to essentially atmospheric pressure within seconds after the event. This is an important consideration in fission product leakage. Tests made with xenon and krypton gas samples indicate the system is effective for retaining fission products in the pool. II-2 ,--g-y-,-- 7,-. + - e u w- -m +--*--r--- --=*--+~e +emmT--+------m^-

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III. C'NDENSING TESTS A. Purpose of Tests The condensing test facility was used to demonstrate and evaluate the condensing effectiveness of steam injection equipment of sufficient size and arrangement that it could be used in multiple units for a full scale pressure suppression system. The specific objectives were as follows : 1. Demonstrate, with steady state steam flow, that rapid and effective condensation may be obtained by a simple straight pipe injector immersed in a pool of water. 2. Investigate the effect on condensation effectiveness of injector parameters such as; injector diameter, steam flow rate, depth of submergence of the discharge, direction of discharge and multiple jets. 3. Evaluate the condensation effectiveness of single and multiple injectors in a restricted volume of water as a function of the geometry of the restricting volume. B. Test Procram The condensing test facility program followed, in general, that initially proposed. The major effort of the program was directed towards obtaining data to evaluate the design parameters of the combined injector-pool arrangements. The sequence of testing followed a logical order of shake-down tests, injector parameter tests, including a series of tests to determine vibrational characteristics of the injectors, and compart-ment tests to determine the interaction between pool geometries and the injectors. Initial Tests The first tests were run with a 4 inch diameter injector for shake-down and instrument check out. These tests were run with an open tank and visual operation of the jet behavior. Initial instrumentation check out and adjustments were made at this time. Group I The first series of tests were to obtain preliminary data for the facility by observing the behavior of steam injected into subcooled water with different flow rates and depths of submergence of the injector. The single 4 inch and 8 inch vertical injectors were used III-1

4 at depths of submergence ranging from 2.1 inches to 2 feet, with steam flow rates of 15,000 to 80,000 lbs. per hour. These tests were run with both the open tank condition (for visual observation) and the closed tank to obtain evaluation data. Group II Before the full scale parametric test program was begun, a series of tests were run to determine the sensitivity of the test facility for detecting steam release into the vapor space. To do this, steam was metered throug/4" diameter orifice was mounted to the h a small orifice directly into the vapor space. A plate with a 3 14 inch by 4 inch reducer section. Attached to the orifice plate and separated by spacer bars, another plate served as a deflector te diffuse the steam into the vapor space. Flow rates up to 500 lbsh.r were used. Group III The tests of Group III were perfortred to determine the condensation effectiveness as a function of the injector parameters in a large water pool with steady state steam flow. The parameters include the diameter of the injectors, steam flow rates, depths of submergence, and direction of jet discharge. The nozzles used for these tests were the 4, 6, 8 and 14 inch diameter single vertical injectors, a multiple 4 inch diameter injector, a 4 inch diameter top horizon al 4 injector and 4, 6, 8 and 14 inch diameter bottom horizontal injecters. All of the injectors are shown in Figure III-4. During these tests, the tank was both open (for visual operation) and closed to determine condensation ef fectiveness. Depths of submergence ranged from 6 inches to 6 feet with steam flow rates of 10,000 to 93,000 pounds per hour. Group IV i During the tests of Group III it was noted that severe tank vibrations occurred with some of the test conditions. A series of i tests was performed to provide some insight into the magnitude and frequency of the pressure fluctuations that caused the vibration. i For these tests, the pool temperatures were varied from 50'F to 150'F and the flow rates were varied throughout the maximum range. In addition, the depths of submergence and support of the injector proper were investigated. The tests were made using the 4, 6, 8 and 14 inch diameter single injectors and the multiple 4 inch diameter injector. i III-2

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Group V To evaluate the condensing effectiveness of the various injectors in a confined volume of water, comparable to that available to individual injectors in a prototype, a series of tests were run with the injectors discharging into a confined compartment that could be varied in width, length, and depth. It was also possible to have various combinations of the width, depth and length. The 4" and 8" diameter vertical injectors and a multiple 4 inch diameter injector were used for these tests. The multiple injector was mounted in both the tandem and side-by-side positions. The tandem position is shown in Figure III-3. The side-by-side position was with the compartment 18" wide and the line of injectors parallel to 4 the width of the compartment. The 4" and 8" injectors were used with 6" and 12" wide compartments, respectively. C. Descriotion of Test Facility The test facility was located at the Moss Landing Power Station of the Pacific Gas and Electric Company. Figures III-1 and III-2 show the general arrangement of the facility. l The basic test chamber was a tank 20 feet in diameter and 24 feet high with dished bottom and top. Figure III-3 is a cut-away isometric view of the facility showing the major details to be described. The top was penetrated by the following: 20" diameter manhole, a 3" vent valve, pressure relief valve, rupture diaphragm, 3 thermocouple leads for measuring temperatures inside the tank, and 4 glass windows for observing the inside of the tank. The side of the tank contained a 6' diameter hatch for equipment, water level gage connections, two 14" diameter steam inlet headers, two 1-1/2" connections for filling the tank, a 4" drain and 14 glass viewing ports. Attached to the exterior wall were a ladder to the top, and two platforms so personnel could look through elevated windows. A standpipe was attached near the bottom of the tank to control normal water level and to provide an overflow line for condensate during testing. r Four flood lights of 500 watts each were used to light the interior of the tank. The interior of the tank was painted white for better visibility. Steam was supplied from two existing 8" - 100 psi steam lines. Headers from these lines connected to a single 10 inch line which contained an orifice for measuring steam flow. An 8 inch angle valve located 5 feet from the base of the tank was used for flow control. After i the valve the line branched to enter the tank at elevations of III-3 _. _.. _.., _ _. _ _.. _ - _ _ _._ _.~._..

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i l' - 6" and 22' - 6". Both lines were 14" diameter. Special combination ring-blind flanges were used in the tee at the bottom to prevent steam flow to the unused header. The 8" control valve was bypassed by a 2 inch line containing a flow control valve and two drain valves. The bypass valve was used for flows under 1000 pcunds per hour. Instrumentation for the compartment tests consisted of windows for visual observation and static pressure taps on the side of the compartment, and one thermocouple located just below the water surface (supported by a wood float) at the end of the compartment away from the injector. The static taps were used to measure surface deflections. Both still-sequence pictures and movie pictures were used to record the pool and jet behavior within the compartment. Basic instrumentation was that necessary to measure the pressures, temperatures, water level and steam flow. Figure III-3a shows the instrumentation line diagram. Steam flow into the tank was measured in the line before the flow orifice and in the 14 inch line just before it entered the tank. Steam pressure downstream from the control valve was also measured. After testing had begun instrumen-tation was added to measure steam pressure at the discharge end of the steam injector pipes. Vapor space instrumentation was for the purpose of detecting incomplete condensation of steam with a closed tank. A water level gage was used to determine the vapor volume above the water pool, 3 thermo-couples were used to measure the vapor space temperature, and the pressure in the vapor space was measured with a mercury manometer. The group II runs were made to determine the sensitivity of this means of detecting steam release. However, it was subsequently established that small amounts of steam released to the vapor space when the tank was open could be readily detected because of fog formation. Most of the tests were run with the manhole at the top of the tank open for visual observation. Water temperatures within the tank were measured with 3 thermocouples located in the side of the tank. All 6 thermocouple temperatures were recorded by a 16 point per/ minute Speedomax Recorder, with repeating preference given to those considered most important for a particular test. For the tests with a confined pool volume, a compartment was fabricated with over-all dimensions of 12' long,12' deep and 18" wide. The compartment was so arranged that the following variations in dimensions could be obtained: III-4

Width : 6, 12, 18 inches 1.ength : 4, 6, 8,10,12 feet Depth : 6, 8,10,12 feet Any combination of these dimensions could be used. Glass view windows were located on the sides of the compartment to observe the surface of the pool near the injectors and the steam jet at the injector discharge. The compartment was positioned and supported on a frame that could be relocated easily within the tank. Test Procedure The standard practices for operating this type of experiment, such as tank filling, steam warm-up, and instrument checks, were followed in preparing for testing. A minimum of three people were required to run the tests. The operator manipulated the steam control valve, filled and drained the tank and warmed up'the system for testing. In addition, he was occasionally called upon to record data. The recorder transcribed the data from the instruments to a log at specific time intervals and directed the operator's activities. The observer maintained positions on top of the tank, on the ladder, or on the platforms and recorded the reactions within the tank. For the compartment tests, the operator was requested to spin the control valve open quickly and set the required flow for the duration of the run. Some of the runs were as short as 20 seconds. After each run, cold water was pumped into the compartment with a fire hose to prepare for the next run. Most of the runs with the compartment were recorded with movie and still-sequence pictures. The facility was secured in the manner requested by plant officials as would befit any facility of this type. C. Discussion of Results Group I These tests were performed to observe the effect of large flow rates of steam being injected into a pool of subcooled water, and to set the practical operating limits of the facility. The initial test was carried out with the single 4 inch vertical injector. Depths of submergence 4 i used were 6 inches and 1 foot. Steam mass flow rates to a l maximum of 53,300 lbs/hr were injected. Jet flow through l the injector was observed at 14,200 lbs/hr. There was no evidence of release of steam from the pool during the test at either depth of submergence or at any of the mass flow (' rates. The preceding operation was repeated with the 8 inch injector at mass flow rates to 90,000 lbs/hr with j the same results. m. III-5 1

From the results obtained it was concluded that the facility was adequate and needed no further modification. Some of the highlights of the visual observations provide interesting aspects in further describing the reactions in the pool. During the operation of the single 4 inch vertical injector-in the open tank run, the pool surface was carefully observed. As the mass flow rate increased, the agitation of the pool increased. At 30,000 lbs/hr (recorded) small vortices appeared around the injector for approximately 10 seconds and then disappeared. A few seconds after the small vortices disappeared one large vortex began forming around the injector and as best one could see extended just below the injector. This lasted abcut 10 to 15 seconds followed by a tremendous upheaval around the injector. This, of course, was the release of the air that had been drawn down the vortex. With the 8 inch vertical injector installed and a mass flow rate of 70,000 lbshr only the small vortices around the injector appeared. These collapsed but never formed into one large vortex. Groun II These tests were performed with a 3/4 inch orifice (See Figure III-4) installed on the 14 inch header and a sealed tank. To establish condensing effectiveness it is important ~ to be able to detect and if necessary measure, any significant steam release to the vapor space above the pool. Part of any steam released would be condensed on the tank walls and top, and on the pool surface. Results of the test using the orifice to meter steam directly to the vapor space show that the rate of condensation is negligible (less than 500lbs/hr). Calculations show that any significant amounts steam released to the vapor space will cause rapid and substantial increases in both pressure and temperature. 1000 lbs/hr will cause the pressure to rise about 1/2 psi 0 per minute and the temperature about 13-1/2 F per minute. The temperature chart for this run is shown in Figure V A-9. Group III These tests were performed to determine the relationship between injector geometry and condensing effectiveness. The tank was open during the first tests of each injector and then closed for the remaining tests. The first tests were for visual observations of the reaction at various flow rates and depths of submergence. The remainder of the tests, during which the tank was closed, provided data that con-curred with the visual observations indicating whether or not steam was released. The only indication of steam III-6

release occurred during the tests with the 6 inch and 6 inch' bottom horizontal injectors at 6" depth of submergen:e. This could be visually observed during the tests at the higher mass flow rates when the momentum of the jet leaving the injector was so great that it carried across the tank befcre completely condensing and the splashing against the wall of the tank released some of the uncondensed steam. With the 6 inch single vertical injector vortices first appeared at a mass flow rate of 48,400 lbs/hr. The vortex lasted about 10-15 seconds and then collapsed followed by the upheaval of air through the water. The mass flow rates were held for about one minute but the vortex did not referm. At each 5,000 lbs/hr mass flow rate increase vortices formed and collapsed. The vortices were not self-sustaining at any mass flow rate. At a maximum mass flow rate of 83,300 lbs/hr, agitation of the pool was severe and the direction of flow of the pool surface was from the ihr wall to the injector indicating the type of circulation apparent in the pool. During a run with the 4 inch single vertical injector, rate of 58,000 lbs/pth of submergence and maximum mass flo closed tank, 6" de hr, the tank was quiet during the first 16 minutes of operation. At this point, the pressure in the tank rose to about 6" mercury due to the accumulated condensate and the tank commenced shaking. Two minutes later, the tank began to shuddir and then bang severely. This banging literally sounded 'like rapid fire from a ri fle. After 5 minutes of this the vibration was severe enough to shake open the safety relief valve and the run was secured. It was concluded from this run tha t the pool vibration phenomenon was temperature sensitive and that the frequency and magnitude of the vibration should be investigated. The 14 inch injector was too large for the steam supply available. At the low line pressure it was not possible t: obtain jet flow with this injector and chugging occurred. With the tank closed, shaking and banging occurred but the intensity was less severe than that of the smaller injectcrs with higher velocities. The 4 inch multiple injector provided the smoothest running of all injectors tested. Tge tank bagan to shake as the pool temperature neared 150 F but the intensity was very low. The 8 inch injector reacted essentially the same as the other injectors in causing shaking and banging in the tank. During the second run of this series at a mass flow rate of 77,500 lbs/hr, the water level was dropped, via pumping, to observe at what depth of submergence steam would be released. Steam blow by was observed at 1" depth of submergence. III-7

The 4 inch top horizontal injector ran smooth during the runs at lower pool temperature with 6" depth of submergen:e. At 6' depth of submergence it became somewhat rougher as the water temperature increased. The bottom 4 inch injector operated the same as the previously run top horizontal injector. The operation was smooth at 6" depths of submergence and gradually increased in roughness at 6' depth of submergence as the pool water temperature increased. The operation of the bottom 14" injector proved to be a serious problem since the steam line ran horizontally from the control valve into the tank. With hot steam on one side of the valve and cold water on the other, along with insufficient pressure to produce jet flow, the water hammer was quite severe. The tank shook severely during the run which was made at 6 inch depth of submergence with the tank open. As the steam valve was being closed, at a slow rate, the severity of the water hammer increased to a point where everything in the vicinity of the tank was shaking. Grouo IV The 4, 6, 8 and multiple 4 inch injectors were each retested. This series consisted of tests with each injector rigidly connected to the tank wall at a point midway over the length of the injector and corresponding tests with the rigid connection removed. These tests were run to investigate the ef fects of various parameters on the magnitude and frequency of the pool pressute fluctuations generated by the discharge of steam. An oscilloscope with a camera was used to record the output of a pressure transducer mounted in the tank wall at approximately a 5' elevation in line with the injector. In each case, the runs consisted of going from maximum mass flow rate to zero in increments of 15,000 lbs/hr in a cold pool at 6' depth of submergence. Then at maxin:um mass flow rates, the pool was heated. After maximum pool temperature was reached the mass flow rate was reduced by ( 15,000lbs/hrincrements. The depth of submergence was then lowered to 4 '. Mass flow rates of 105,000 lbs/hrand75,000 lbs/hrrecordedwereused. The depth of submergence was lowered to 2' and the process repeated. Photographs from 0 the scope were taken at each mass flow rate and at each 10 F temperature rise during the heating run. Figures V A-1 throug V A-8 are a complete series of photographs taken during the test of the 8 inch vertical injector anchored and unanchored. III-8

During the operation of the 6" injector anchored to the tank wall and specifically during the run where the mass flow rate was being reduced following the heating of the pool, the facility became so rough that the tank appeared to be bouncing on its foundation. Personnel in the control room of the power station reccrded a mild earthquake which time-wide coincided with this run. Photographs taken during this run show peak pressure fluctuations of 6 psi, while photographs taken when the anchor to the tank wall was removed show peak pressure fluctuations of 10 psi during the same run with less severe reaction of the facility and no bouncing of the tank. In this same unanchored series, at 4' depth of submergence and recorded mass flow rate of 105,000lbs/hr, the peak pressure was 24 psi as compared to 3 psi when anchored. At 2' depth of sub:ergence a peak pressure of 40 psi + as compared to 6 psi when the injector was anchored. It was observed during the entire series of tests in thig group, that as the pool temperature reached 120 F to 130 F the roughness commenced and increased in intensity with the increase in temperature. The operation of the facility was smooth in all cases as long as the pool temperature remained below 120"F. l l' III-9 l-

Compartment Test Results Speaking generally of the compartment tests, the pool surface-i within the confines of the compartment walls was much more agitated than had previously been observed with the same injec. tors i } discharging into the large pool. There was no tendency to form vortices, but the surface in the neighborhood of the injector tended to be depressed below the surroundings. A typical example of this behavior is pictured in Figures III-6 and III-7, which are top and side views, respectively, for run No. 54 (the first three 4: pictures in each figure were taken before the steady state pattern was established). The surface typically had a very ' foamy appear-l ance, particularly at the injector end of the compartment and at the end opposite the injector. Water flowed down from the end opposite the injector toward the injector end; there was a small jump just before it reached the injector. l The viewing windows in the sides of the compartment permitted a relatively clear view of the jets which fermed at the injector discharge. For each run these jets underwent a characteristically j-changing pattern. Some of the characteristics are shown in Figure i III-5. Picture 1: No flow. Picture 2: The first puff of steam appears. Pictures 3 and 4: Full flow but the jet is very short 4 (in this case, about 4" to where it " necks down". The calculated i. velocity for this picture was 535 ft. per sec.), associated with high subcooling. Pictures 5 and 6: The jet becomes longer and is ] less " necked down" as the pool subcooling decreases. Picture 7: Apparently slightly underexpanded jet, low subcooling (calculated velocity for this picture was 1000 f t, per sec.). The jets some-i times tended to draw together as shown here. Picture 8: Shutting down. Note in the last four pictures the formation of air bubbles in the bulk of the pool to the left of the jets. These appeared as the water temperature was approaching 150*Fi. They are presumed I to be initially dissolved air which had been driven out of solution. i The results of the compartment tests are sununarized in Table I, s l (except for qualitative observations of the effect of air on con-densation, which will be described later). Columns are arranged r to give, for each run, geometry, steam flow rate, and two quantita-tive measures of performance, viz., surface deflection, and mean temperature rise (calculated) versus temperature rise at a point (measured). The following notes are intended to better identify some of the quantities. l 1. " Injector - Distance From End" is the distance from the end of the compartment to the center of the injector (to the center of the nearest nozzle in the case of the triple injector). j 2. " Rates - Mass" is the Barton Flow Meter reading corrected for steam conditions different than rated. For runs 84 through 91 ) the single 4 inch injector was itself calibrated (the minimum flow which can be measured satisfactority with the Barton Meter was too great for these runs) to give mass rate as a function ofop across the injector and the state of the steam directly upstream. f 111-10 p.

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TABLE III-I COMPARINENT TEST RESULTS Compartment In1ector Rates Deflection Temp. Rise Sub-Diet CALC. Run Length Depth Width merge From Size

o. Mass Velocity MOM.

MAN. OBS. MEAN.N!AS. (1b) (ft) (ft) (OF) (OF) No. (ft) (ft) (ft) (ft) Ind(ft)

1b/sec)

(ft/sec) Triple 1 12 12 1.5 6 0.7 4" 17.1 1470 775 51 52 18.2 1510 840 3.0 39 49 2 3 20.7 1580 975 64 49 4 23.7 1610 1115 66 40 5 10 12.4 1310 470 21 30 6 14.4 1340 600 53 31 7 16.7 1450 755 76 52 8 10 / U ::0 C/" e-9 16.7 1530 830 75 59 10 20.2 1560 945 66 56 gg 21.9 1590 1040 > 1.2 76 56 12 4 7.8 785 190+ 0 46 42 }} 10.3 1035 3301 0 52 25 14 12.8 1230 490 46 30 15 16 17 8 11.2 1110 385 0.6 64 58 18 13.3 1265 530 0.8 58 36 19 16.6 1450 745 1.1 64 64 20 18.3 1510 850 1.4 63 65 21 20.4 1570 960 1.7 79 63 22 10 13.3 1265 530 1.8 91 86 23 18.6 1525 895 67 40 24 25 12 9.7 965 290 0.1 29 33 26 13.5 1280 535 0.3 59 54 27 19.5 1550 915 0.3 54 38 28 8 9.4 935 275 0.1 36 48 29 13.4 1275 535 0.2 44 46 30 19.0 1535 885 0.4 59 32 31 8 8 1.5 6 0.7 10.3 965 310 0 81 0 32 14.8 1230 565 0.1 150 6 33 20.1 1420 885 0 137 3 34 12 12 0.5 35 5.3 530 87.5 0.5 0.2 44 45 f

Table III-I (Continued) Compartment Iniector Rates Deflection Temp. Rise Sub-Dist CALC Run Length Depth Width merge From Size Mass Velocity MOM. MAN. OBS.MEAN MEAS. No. (ft) (ft) (dt) (ft) End(ft) [1b/sec) (ft/sec) (1b) (ft) (ft) ( F) ( F) Triple 36 12 12 0.5 6 0.7 4" 10.2 1110 355 5.0 1.2 67 42 37 15.0 1425 665 7.3 72 27 38 8 5.3 540 89.5 0.6 0.2 47 34 39 10.3 1155 370 3.0 0 77 39 40 14.6 1435 650 4.2 118 42 41 8 5.3 530 87.0 161 78 42 Triple 43 12 12 1.5 4" 5.3 525 56.5 0 0.2 13 22 44 side 10.2 1005 320 0.2 25 34 45 by s ide 14.6 1320 600 1.0 1.2 40 25 46 8 5.3 520 85.5 0.1 0.3 22 25 47 10.2 1000 315 1.0 0.5 28 32 48 14.6 1325 600 1.4 2.2 49 12 8 5.3 515 85.0 0.1 0 20 26 50 10.2 995 315 0.3 0.4 31 32 51 14.6 955 435 1.0 0.5 64 44 52 8 5.3 495 81.5 0 0.2 34 44 53 10.2 975 3 10 0.7 0.8 35 44 54 Triple 14.4 1295 ! 580 1.6 1.9 100 107 55 12 12 0.5 4" 5.4 535 '50.5 0.6 0 58 43 56 7.8 805 195 1.5 0.3 80 60 57 10.2 1000 320 1.6 1.5 74 40 58 12.5 1210 475 6.1 2.5 h, 59 14.4 840 375 4.2 66 65 60 8 5.4 535 90.5 1.0 0 158 76 61 7.8 715 175 1.3 0.4 199 75 62 10.2 950 300 2.3 2.5 63 12.5 1230 480 3.0 2.5 136 72 64 14.7 1260 575 3.7 2.5 236 75 65 5.3 515 85.5 1.8 0 44 67 66 7.9 770 190 2.1 5?8 68 77 67 10.2 835 265 5.2 [35 84 101 68 12.6 1030 400 7.5 2.0 129 97 69 8 I 5.3 545 90.5 1.7 0 69 91 70 7.8 740 180 3.3 5?S tti 101 fb 80 78 71 10.3 1065 340 1.7 72 12.6 1045 410 4.0 270 260 62 73 12 12 3.2 5.3 540 89.5 1.0 g,0 74 10.2 1075 340 2.5 0.5 52 25 75 14.8 1040 475 3.0 0 133 50

Tcble III-I (Continued) Compartment Injector Rates Deflection Temp. Rise i Sub Dist

CALC, Run Length Depth Width merge From Size Mass Velocity MOM.

MAN. 03S MEAN MEAS. No. (ft) (ft) (ft) (ft) End(ft) (1b/see) (ft/sec) (1b) (ft) (ft) (CF) ( F) 76 8 8 8 1.0 4 0.7

SE*

5.3 415 68.5 0.2 0 55 47 79 10.3 705 '225 1.9 37$* 54 58 80 14.8 990 455 3.1 2Vgr 80 68 81 12 12 6 5.3 390 64.5 0.4 0 18 27 82 10.3 740 240 0.6 0 19 22 83 15.2 1030 485 1.9 {;@ 30 16 84 6 6 0.5 3 single 85 1.1 340 11.5 -0.4 0 86 3.3 810 81.5 -0.8 0 87 5.8 1030 190 1.5 0 190 38 88 12 12 6 89 90 3.0 755 70.5 0.4 0.5 11 14 91 5.1 1055 173 O.8 2.0 29 15 1 4 4 1 b

" Rates - Velocity" is the calculated velocity at injector exit. " Rates - Momentum" is the product of mass rate and velocity. The bases for these two calculations are given in Appendix B. The method of calculation could not be applied to runs 1 to 30, because no pressure at injector exit was measured. Velocity as a smoothed function of mass rate was determined, based on all the subsequent runs and then applied to runs 1 to 30. 3. " Deflection" refers to pool surface deflection in the com-partment. It is the amount that the surface in the region of the injector was depressed below the surface at its highest point. This proved difficult to determine by direct observatica. For run 35 and after, another indication of this surface deflection was obtained by measuring the difference between the static pressure below the surface at the injector end of the compartment and the static pressure below the surface at the end opposite the injector. The measuring instrument was~a manometer, hence, the column head-ing assigned to this quantity is " Man." Values determined by direct observation are in the column headed "Obs." 4. " Temperature Rise - Calc. Mean" is the calculated temperature rise for the water in the compartment from a consideration of 4 t heat capacity and net rate of energy addition. The basis for I[, I this calculation is given in Appendix B. " Temperature Rise - Maasured" is' the teciperature rise measured at a point located just below the surface and at the end of the compartment opposite the injector. Visual means were employed to check for steam release during the compartment test. No release of' injector steam was detected for any of the runs, although so m fog formation was observed near the end of some cf the runs. In these cap s the compartment water

7

~ temperature was approaching boiling. The duration of the comp artment runs aried frorc about 20 seconds ) (Run #85) to about 100 acconds (Kun #2).. Although a few seconds / were allowed after the beginning of each cun'for,the instruments to come to equilibrium, tie. data is undoubtedly af fected by the time response characteristics of the instrumen.ts, particularly for the shorter runs. The pressure regulating system on the steam supply itself has a relatively slow time response (of the order of / 15 seconds) which further affected the data. The compartment flow s pattern, temperature, and volume occupied by vapor (associated with d' decreasing subcooling), were all in a transition state, further complicating the problem of securing (and interpreting) data. The final effect of the response characteristics of instruments and pressure regulation, nadithe transient character of the compartment phenomena was a scattering of data which carried over into the results in Table 3 7 Any cumulative effects are impossible of detection; it,is' believed that they are minor. t III~I5 m / v ~f / / ,J' A-

s During some of the runs, a general receding of the surface in the compartment, of as cuch as a fuot at the highest flow rates, was noted. This must have been due to the region of high pressure i directly below the injector discharge plus the presence of holes and cracks through which part of the compartment contents could leak out. This outflow at the bottom of the compartment could alter the flow pattern, and certainly would reduce the compart-ment heat espacity. Both these effects are considered small and have not been taken inco account in reducing the data. They probably contribute a seall 4anount to the scattering of the results already referred to. The surface deflection as indicated by the manometer reading is usually not in agreement wi.h (though in general it shows the same trends as) the directly observed surface deflection. The static pressure sensed by the instrument and the hydrostatic pressure corresponding to the surface elevation would not be expected to be equal, _because of the presence of a velocity head and because of the rotational, turbulent, highly dissipative character of the flow. It was anticipated that the relationship between sensed static pressure and surface elevation might be very nearly the same at one end of the compartment as at the other (i.e., the velocity, hence, the velocity head, should be of the same magnitude, etc.) If this were so, the manometer readings and the direct obsersations would agree. It appears, therefore, that the relation-ship between sensed static pressure and surface elevation are not the same at the two ends of the compartment. However, some of the lack of agreement can be attributed to the scattering (already mentioned) which all the data was subjected to in greater or lesser degree. The relative agreement tends to be better at greater deflections. At all deflections greater than about 3 feet the manometer deflection exceeds the directly observed deflection, a happy trait from a safeguards point of view. More than this, the manometer deflection correlates much better with the mass and momentum rates. For the foregoing reasons, the manometer deflection is taken as the measure of surface deflection in the subsequent discussion of results. It will be noted that for a few of the runs a part or all of the results are missing. This is simply due to the fact that the corresponding data were obviously of doubtful accuracy, or the data were obtained under uncontrolled conditions (e.g., runs 15 and 16, during which the baffle forming one end of the compartment broke loose). The effect of air on the complete condensation of the steam which would follow it, under cenditions which would exist in a complete and functioning pressure suppression system, was investigated j qualitatively. The 14" pipe between the steam valve and the injector (see Figure III-3 and description of Condensing Tank Facility) was initially filled with air (Runs 76 and 77). The steam valve was 111-12 L J

m ~ opened as quickly as possible to a set flow, and the manner in which the air escaped from the pool was observed, visually, with still pictures (see Figure III-8), and with movir.g pictures (Reci #4, 64 frames /sec.) These observations demonstrated that all the air, except for some very small bubbles (estimated at less than 1/4 inch diameter) escaped.to the surf ace in of the order of 1/10 second. The time was too short for the bulk of the pool water to be set into motion. The air broke the surface directly above the injector exit. ( What was possible to observe of the behavior of the pool in the Transient Test Facility, both visually and with 64 frame-per-second movies, fitted this same general pattern. No circulating motion was observable just prior to air breaking through except for a slight heaving of the surf ace. The violence of the action upon breaking the surface then caused the view to be obscured, making further observation of air escape impossible). During some tests run earlier and reported in the Phase I Report, stea= containing small amounts of air was discharged into a pool of water. Here the air coalesced at many points to form small bubbles, which then approached the surf ace relatively slowly. Discussion of Compartment Test Results One criterion for the adequacy of any pool-injector combination is the complete condensation of any steam discharged into the pool. The com-partment test was to establish the range of parameters for which this criterion would be met. Relative to the effect of air on the completeness of condensation, it may be stated that any air being driven in large quantities from the dry well and injected into the pool will escape rapidly from the pool, so it cannot affect the mechanism of condensation except for a brief inter-val of time. The Transient Test Facility containment pressure data reveal no evidence of uncondensed steam. On the other hand, air injec ed into the pool in small quantities will become dispersed as small bubbles which will have time to reach thermodynamic equilibrium before breaking the surface. Hence they can carry no more vapor than that carried by saturated air. No more can be said on the basis of these test results concerning the effect of air on condensation. When the iajector is discharging steam into the pool the pool surface is disturbed. If it should be depressed in the region of the injector by an amount approximately equal to the original injector submergence, steam may be released. One necessary condition for meeting this criterion for adequacy, then, is that the pool surface nust never he depressed below the iniector exit. If the subcooling drops to near zero along any possible steam bubble i path to the surface, steam may be released. Let the highest-tempera-ture at which complete condensation is assured be Tc. A second con-dition for meeting the criterion of adequacy is: there must be no relatively continuous regions extending from the injector exit tjg the surf ace of the pool, at temperatures above Tc. 111-13 ___._y c. ,-e.

1 The Table I summary of the compartment test results may now be examined in light of the two conditions just stated. First, evidence is desired to show that the flow pattern is a reliable two dimensional representation of a more extensive system. A comparison of the triple 4 injector arranged side by side in an 18" wide compart-ment, with the single 4" injector in a 6" wide compartment provides such evidence. Surface deflection A z is plotted versus mass rate per foot of compartment width for runs 43, 44, and 45 (triple 4" injector,12' x 12' compartment) and for runs 90 and 91 (single 4" injector, 12' x 12' compartment). The plot is shown in Figure III-9. (As a matter of interest the data obtained by direct observation are also plotted). Surface deflection is plotted versus mass rate in Figure 111-10 for the triple 4" injector in the 6" wide compartment. The depth of submergence is 6' in every case. From a consideration of the first condition for adequacy (i.e., surface depression must be less than injector submergence) the 12' x 12' compartment with injector 3.2 feet from the end is best. An extrapolation would give approximately 50 lbs/see f t. before surf ace deflection would equal depth of submergence. The 8'L. x 12'D. is next (42 lbs/sec ft before 6 z. depth of submergence), the 8' x 8' next (40 lbs/sec f t before A z a depth of submergence), the 12' x 12' next (27.5 lbs/dec ft before At = depth of submergence), and the 12'L. x 8'D. last (22 lbs/sec ft before A z = depth of submergence). It should be pointed 4 out that in the condensing test facility, surface defic-tion and surface depression are equal (except for a few cases where the whole surface was depressed as mentioned earlier). In an actual system, however, no water could spill over the walls of the compartment so that for every depression of the surface in one region there would be an elevation of the surface in one or more other regions. Surface depression would always be less than surface deflection as defined here. However, for the purpose of comparing compartments, surface deflection is as satisf actory as surf ace - depression would be. It should also be recalled that surface deflection l (as used in the above discussion) is based upon surface deflection as indicated by the difference in static pressures measured at the two ends of the compartment. The agreement between observed surface deflection and surf ace deflection as determined above is usually relatively good at high mass rates. For example, refer to Table III-I, run numbers 58 and

59. The " observed" surface deflection was known to be between 2.5 and 5.0 feet (the surf ace was actually below the top window and above the bottom window, hence could not be observed directly), for both runs.

[ The indicated surf ace deflection for run number 58 was 6.1 feet, slightly outside, but on the conservative side of the Ibaits determined by obser-I l vation. The indicated surf ace deflection for run number 59 was 4.2 feet, which is nicely bracketed by the limits set by observation. The agree-ment is worse in some cases and better in others, though the general tendency is for the indicated deflection to be greater than the observed deflection, and therefore conservative, at the higher flows. ) III-14 l l l l \\ ...n. .-,,,.n, -._,,,---.,r

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t t.y h I b FIGURE III-10 1

l It may be argued that good agreement between calculated mean tempera-ture rise and measured temperature rise at a point is evidence that the mixing is good. It is not conclusive evidence, but provides at least a basis for comparing compartments. Good mixing is insurance that the second condition for adequacy (no relatively continuous regions at temperatures above Tc) can be met. The ratio of measured temperature rise to calculated mean is plotted versus mass rate in' Figure III-11 for the triple 4" injector in the 6" wide compartment. The ratio is positive for the 12'L. x 8'D. compartment for most flow conditions, correspond-ing to a higher measured value than the calculated mean. There must have been some streaming action which caused a high temperature stream to flow past the thermocouple used for temperature rise measurement. This is a situation to be avoided per the second condition for meeting the criterion of adequacy, particularly in view of the thermocoupit!s loca-tion near the surface. The ratio of measured to calculated mean tesp-erature rise for the remainder of the compartment sizes, except for the 8' x 8' compartment at flows less than 15 lb/see per foot, is less than unity. There appears to be little to choose between these remain-ing compartments. At flows of 20 lbs/sec per foot of width the ratio is approximately 0.5. The ratio increases consistently as the flow decreases toward zero, and in the case of the 12' x 12' compartment it actually appears to be approaching unity at zero flow. Speaking generally of Figure III-11, there are a few points which are wild and must be questioned, but for the remainder there is a reassuring consistency which argues that mixing is much better at lower flows. It is to be noted that the time when good mixing is important is when the flow is low. This would be the situation after peak response in an actual pres-sure suppression system, when the steady absorption rate of energy would be causing the pool temperature to increase. Surface deflection is plotted versus mass rate per unit width in Figure 111-12 for the single 4" injector, the single 8" injector, and the triple 4" injector in side by side arrangement. The depth of submerg-ence for the single injectors is just half the compartment depth; for the triple injector the depth of submergence is 6 feet. The results for the single 4" injector' in the 6' x 6' compartment are at variance with the rest of the data and with what shocid Be expected. Neglecting the 6' x 6' compartment data, the remainder of the data is relatively consistant. Although there is not enough data to conclusively establish l which compartment is best (the mass rates were much too low to depress i the surf ace close to the injector depth of submergence), the 12' x 12' l with any one of the three injectors tends to be better than the average, l whereas the 8' x 8' with either the single 8" or the triple 4" injector l tends to be worse than the average. The ratio of measured temperature rise to calculated mean for the in-l jector compartment combinations of the last paragraph are plotted versus l mass rate in Figure III-13. Evidence of the streaming action, observed for the triple 4" injector in the 8" deep compartments, is seen for every injector-compartment combination here except the single 8" in the 8' x 8' compartment. (The single point for the single 4" in the 6' x 6' l compartment is inconclusive). It is particularly evident at low flow l in the compartments of 12 foot length, t III-15

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=. Apparently the triple injector (nozzles in line with the compartment) in the 12' deep compartments gives much better mixing than do any of the single injectors (or than the triple injector with nozzles arranged side by side, which is the two dimensional equivalent of a single injector) in the same compartments. It must be pointed out, however, that the triple injector was not tested at as low values of mass rate-per unit width. It has not been established that the triple injector gives superior mixing at very low flows. At low flows in the 8' deep compartments the single injectors appear superior. Conclusions From Compartment Test Results From the results of the compartment tests as described in the preceedigg section, the following conclusions may be drawn: 1. Air driven from the dry well at high mass rates, as would occur in an actual system early in the series of events following a maximum credible type accident, and injected into the pool, will escape rapidly from the pool (in about 1/10 second for injector submergence of 6 feet), so cannot affect the mechanism of con-densation except for a brief interval of time. 2. Low mass rates of air injected into the pool simultaneously with steam, as would occur in an actual system immediately after the initial peak response, will become dispersed as small bubbles which will have time to reach thermodynamic equilibrium before escaping to the pool surface. 3. The two dimensional approximation of an actual injector-pool come bination may be simulated in a compartment having a width of 6 inches, within the accuracy of the data reported here. 4. From a consideration of the first condition for adequacy (surface depression less than injector submergence), for the triple 4 inch injector at relatively high mass rates (greater than 15 lb/see per foot of compartment width, say), the 12' x 12' compartment with injector 3.2' from the end is best; the 8' x 8' next; the 12' x 12' next; and the 12'L. x 8"D. compartment is poorest. (The single injectors were not tested in this range of mass flow rates, so no l conclusions can be drawn concerning their performance relative to i the triple injector.) At relatively low mass rates (less than l 15 lb/see per foot) all injector-compartment combinations gave l comparable (and adequate) performance, sithough the surface deflec-l tion with the 8' x 8' compartment tended to be greater than with the others. 5. From a consideration of the second condition for adequacy (no relatively continuous regions at temperatures above Tc), with particular concern for low mass rate conditions, and on the basis that the temperature measured at a point near the surf ace should be in good agreement with the calculated mean, the triple 4 inch l injector with the 12' x 12' compartment is best, with the 8'L.x12'D. hext best, with the 12' x 12' and located 3.2' from the end some-what uncertain, and with the 8' deep compartments the poorest. III-16

F On the same basis all the single injector-compartment combinations were poor compared with the triple injector. (The foregoing conclusions concerning uniformity of temperature are only tentative for the following reasons: (a) good agreement between temperature measured at a point and calculated mean temperature is not conclusive evidence that the temperature is everywhere uniform. In particular, it does not insure that there are not relatively continuous regions, removed from the thermo-couple location, which are at temperatures higher than the mean; (b) the temperature measured at one point near the surface may not. be representative of the entire surface. In particular, continuous regions of high temperature water whose possible existence was admitted in (a) could extend to the surface; (c) it is particularly important that the temperature be uniform at low flows. No. data was obtained at flows less than 10 lb/see per foot for the triple injector, or 3 lb/sec per foot for the single injectors). III-17

4 IV. 'IBAIISIEllT 'IESTS A. Purpose of Tests he primary purpose of the transient test facility was to demonstrate, with a working scale model, the effectiveness of the pressure suppression to contain and absorb the energy and radioactive materials released from the rupture of a reactor pressure vessel or associated primary piping. To this end the specific objectives were as follows: 1. To obtain transient pressure data as functions of the major system parameters that'would confirm the analytical model' or to provide the basis for improving it. 2, 2 provide pressure response data not specifically included in the analytical model such as; the pressure decay behavior of the dry well and the containment volume. 3 To evaluate the effectiveness of the system to trap and retain fission products that could be released from rupture or meltdown of the fuel elements. B. Test Program he test program followed, in general, that proposed in the Phase I report. The major deviations were to acconnodate imposed test conditions and to expand certain areas of test data. Se actual test-schedule is given in Table IV-1. Most of the early tests were run and rerun to establish satis- ) factory methods of rupturing the discs and maintaining saturated ~ conditions at the bottom of the pressure vessel. Rese tests also were used to determine the necessity of additional instru-mentation such as dry well and containrent volume thermocouples and the differential pressure transducer between the dry well and containment volume. l A comparison of the proposed test pregram and the actual program i indicate the deletion of several test points. Most of these tests were those to be run with different amounts of water in the pressure vessel. It was concluded that the initial amount i of water charge had little bearing on the peak pressure response. More important, it appeared desirable to have a minimum charge of water, corresponding to approximately 80% moisture, to maintain the circulation up to 1000 psi. Other points that were omitted ~ were those that would have the least influence on the program. These were intermediate points for the pressurized dry well, an intermediate dry well condition, and additional points with small orifice openings. l ll-IV-1 i . _. _ _.,. _. _ _.. ~ _ _.

TABLE IV.1 SAN 8IENT EST FACILITY _EST SCEDUIE Point Pm ssung Inital Break Yolums Vent Vent 2 3 2 No. Date psi Moisture Area ft Dry Well ft 3,,,fg Depth ft 6 1 22 1000 .66 .07786 26.00 2.12 15 7 1 27 1000 .66 .o7786 17 53 2.12 15 8 1 29 1000 .66 .00786 17 53 2.12 15 11 1 31 1000 .66 .00177 17 53 9" 15 12 1 30 1000 .66 .o1572 17 53 9" 1.5 13 26 1000 .66 .00T86 17 53 9 75 14 2 1000 .66 .00786 17 53 9(1") o(15)* 16 2 24 1000 .8o .00T86 26.00 9(1") 15 17 25 1000 .66 .01572 26.00 9" 15 19 24 1000 .66 .00786 26.00 .75 i 20 2 19 1000 .80 .00T86 26.00 o(15)* 24 2 18 1000 .8o .o 7786 20 32 9" 15 25 12 1000 .80 .00T86 17 53 9" 15 2") 15 26 2 28 1000 Kr .So .00T86 17 53 9 27 3/2 1000 xe .80 .00786 17 53 9(1-1/2") 15 29 2/11 1000 .66 .00T86 17 53 9(1) 1.0 l 5(l')/2") 30 2 27 1000 .80 .00786 17 53 15 9(1-1 15 31 600 .80 .00786 17 53 32 3 3 1000 D 0 .80 .01572 17 53 9(1-1/2") 15 2 33 3/14 1000 xe .80 .o1'572 17 53 9( 1 ") 15 1000 I+ Ens.80 .01572 17 53 .25 15 34 1000 NaI .So .01572 17 53 .25 15 35

Pressurized Dry Well to effect no water in vent pipes.

l Additional points were included to detemine the effect of a smaller vent area, the effect of an intermediate depth of submergence, and the effect of vent pipe dia eter. The last tests were run using tracers to evaluate fission product entrainment. These tests consisted of runs with xenon, krypten and heavy water as tracers. C. Description of Test Facility The transient test facility consists in essence of three interconnected pressure vessels that simulate the reactor pressure vessel, reactor containment vessel and the pool containment vessel. A simple schematic arrangement of these volumes for a prototype power plant are shown as volumes (1), (2), and (3), respectively, in Figure I-1. Such an arrangement did not permit the accessibility and flexibility necessary for obtaining test points to fulfill a multi-parameter test program. Therefore, the test model was arranged as shown in Figure IV-1, where volu=es (1), (2), and (3) correspond to the reactor pressure vessel, reactor containment vessel and the pool containment vessel, respectively. A picture of the test installation is shown in Figure IV-2. A brief explanation of the operation of the facility will facilitate further discussions of the system. Water in the reactor vessel is heated to saturation conditions at high pressure and then discharged, through an orifice plate into the dry well, by breaking a rupture disc. Part of the water, after it flows through the orifice, flashes to steam. 'Ihe steam-water mixture is then discharged from the dry well, through the vent pipes into the water pool where the steam is condensed. The scale size for the test facility was detemined largely by the size of the pressure vessel used to simulate the reactor vessel. This vessel was in storage at the VBWR site and essentially ready for use. Fortunately, the vessel volume was a factor of one thousand smaller than the proposed 50 My reactor vessel. Using this volume factor and the dimensionless scale factors as determined in the analytical model, the area scale factor was 100 to 1 and_the linear scale factor was 10 to 1. A time scale of 10 to 1 was accepted, rather than distort physical dimensions to accommodate the inability to scale intrinsic fluid dynamic properties. Reference Drawing: 196 E 929 - Arrangement - Pressure Suppression Test Vessel Instmmentation 'Ibe p:imary instrumentation consisted of the pressure transducers and the four channel Sanborn recorder. The pressure transducers on the pressure vessel, dry well, and the containment volume were of the strain gage type made by the Statham Co. They were rated at 1000 psig, 100 psia and 15 psig, respectively. The fourth pressure transducer, used to measure IV-2 -,,e. y y

TRANSIENT PT-1 I NS U L A TIO?.; PRESSUR E RE515ThllC E TEST FACILITY \\ HE h*' I' 5 o PT = PRESSUR E PRESSURE o TenA/soucEe CHIMNEY FOR TC = THERMOCOUPLE V ESSE L T = THERMOMETER c rcutrTicM VOLUME @ - ? TC-l PT-2 w 'gp54 PT-4 (A P) p OR IFIC E ( RU PTUREj DISK O,A B LAST CAP CONTAINMENT \\\\ TC-6 u. VESS E L DRY VOLU M E @ ll : VOLUME WELL 1[ cle-SAMPLE TUBE r-a -VENT l PI P ES POOL y ~ FIGUP" M-N -1 1

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the differential pressure between the dry well and the containment volume vas a 50 psi reluctance type transducer made by Pace Engineering Co. Excitation voltage for the transducers on the pressure vessel and dry well was supplied by 6 volt batteries. Se output signals for these two devices were fed into stabilized D. C. preamplifiers of the Sanborn recorder. Se containanet vessel transducer and the differential pressure transducer were coupled to carrier preamplifiers of the Sanbem acorder. The carrier preamplifiers supplied 2400 cps excitation signal for both transducers. The naponse of the Sanborn recorder is limited tc about 100 cps. Se stabilized D. C. preamp 11fiers have an output Jack with a frequency response of about 10,000 cps. max. Se pmssure transducers all have natural frequencies on the order of 10,000 cps or better. An c cilloscope fitted with a camera vas used in a few of the early tests to check the response characteristics + of the Sanborn recorder. No deviation in response could be detected. A Bounion tube pressure gage was mounted on the pressure vessel for convenient checking of pressures during the heating Period. Temperature for recording test conditions and monitoring during the heating period were measured with chromel-alumel thermocouples and a General Electric potentiometer. The thennocouples vem numbered and located as follows: 1. Located in the pressure vessel 8 inches above the face of the 10" flange. 2. Located in the 6 inch discharge tube app oximately 6 inches above the orifice plate. 3 Lccated in the 6" discharge tube approxi=ately 4 inches above the orifice plate. 4. Located with the blasting cap to monitor the cap temperature during the heating period. 5 Located in the dry well approximately 3 inches down from the 24" flange of the dry well. 6. Located in the dry well approximately 32 inches down from the face of the 24" flange of the dry well. ~l Located in the containment volume approximately 12 inches down from the top head on the northeast side. Se pool water temperature was measured in some of the latter runs with an insnersion mercury thennometer. Se level of the pool was measured with strips of measuring tape strapped to vent tubes that were visible through the sight ports. A water manometer was used to determine the pressure in the dry well for i the tests conducted with a pressurized dry well. A conductivity probe was inserted 0.t the discharge end of a vent tube to detennine the instant the IV-3 --,.m.-- .,m-- -,e-- +v--- -~-, - - - - - - -. -r- +-r

l vater was expelled from the vent tubes. Eis signal was recorded using the external timing trace circuit of the Sanborr. recorder. Reference Drawing: 218 B 890 - Instrumentation Diagram i Pressure Vessel The simulated reactor pressure vessel, shown as the insulated vessel in Figure IV-2, was fabricated from 10 inch schedule 80 pipe, flanged on one end and capped on the other. De vessel is approximately 70 inches long with a volume of thne cubic feet. The discharge passage was obtained with a 6 inch pipe velded to a special reducing-mounting flange to which the 10" flange of the pressure vessel was bolted. Se 6 inch pipe was fitted with the flanges necessary to hold the orifice plates, rupture discs, the shock tube and thrust balancing impact plate. Se shock tube had two extension pieces for use with the tests having diffennt dry well volumes. l Re choice of 6 inches for the discharge and shock tubes was based on i having as large diameter as practical to prevent excessive bubble formation in the discharge section above the orifice and to minimize the back pressure on the discharge side of the orifice due to " choking" in the shock tube. ~ Both of these effects are in the direction of. reducing the mass flow rate. The physical dimensions of 900 pound class flanges made any sizes larger f than 6 inches undesirable. A view of the pressure vessel showing the details of the discharge section is presented in Figure IV-3 The orifice plates were made of 1/2" thick stainless steel with a " sharp j_ edge" of 1/16 inch thickness and a 450 bevel. Bree diameters were made measuring.599 inches, 1.199 inches and 1 599 inches, n spectively. Each of these plates had a dish of approximately 90 mils after the first run which did not appear to become greater. These plates naained in good condition, with respect to the shart edge, throughout the test program with the exception of those used for the fission product containment tests. (Pieces of the glass sample bcttles roughened up the upstream edge during these tests.) The thrust balance plate at the bottom of the shock tube was designed to withstand approximately 70,000 pounds of thrust. Se support bars were turned to a diameter slightly smaller than the thread roots on the plate end and were velded at the top end only with as such fue length as could be conveniently obtained to provide "springyness". Over design of the thrust balance system was felt justifiable to assure the pmssure vessel would not become a missile. The pressure vessel was fitted with a " christmas tree" on the top or capped end to mount the valves and pressun instruments. Rese consisted of a blow down valve, a safety valve, a pressure transducer, a pressum gage, and the temperature bulb for the power controller. In addition, there ven bayonet thermocouple fittings located as follows: one in the side of the pn saure vessel 8" above the face of the 10" flange; and two located in the 6" discharge tube (4" and 6", respectively, above the face of the 6" flange). A 1/4" pipe drain line was located 8" above the face of the 10" flange. IV-4 v~-m, .,,-,.,-.,,_,_.--4 c...--o,--._ ,,-ww,,,-y-,.w,. ,,,,-%,-me,,--, --+w ..-w,...,,,----.eu,,-. ---,, -,, - - - -, - - - - -, -.

i -i%. hgh\\Y?-l4 l o s i l TRANSIE:TT TEST FACILITY Discharge Tube of the Pressure Vessel

An array of formed calrod and strip resistance heating units were mounted on the external surface of the pressure vessel. Rese units were held in place with steel bonds and covered by sheet steel vith 1-1/2" of 85% nagnesia insulation over the sheet steel. An additional calrod heating unit was mounted below the special reducing-mounting flange. De installed heater capacity totaled 33 5 KW; however, due to difficulties experienced in obtaining fully saturated conditions up to 1000 psi only 14.1 KW of input were normally used for heating. This was obtained with the lower bank of 2 calrod units and the calrod unit mounted on the special flange, each rated at 4 7 Kv, 220 volts. Uniform heating of the water was successibily obtained by installing a sheet metal chimney device inside the pressure vessel to promote circulation during the heating period. Bis device consisted of a bottom section 4 inches in diameter joined by a transition section to a top section eight inches in diameter. he bottom section extended within a few inches of the bottom (the orifice plate) of the 6 inch discharge tube. Be top section extended 42 inches into the 10 inch pressure vessel. he top vas ended with a perforated section and a ring cap to improve the circulation pattern. Bis chimney extended approximately 12 inches above the lover bank of calrod heating units. Dry Well and Yent System The dry well volume, number (2) of Figure IV-1, was fabricated from a section of 24 inch pipe joined to a section of 40 inch pipe with a conical transition piece. A quarter top view of this vessel is shown in Figure IV-4. B is unit was mounted on the containment tank by means of a 42 inch by 24 inch reducing flange. The 24 inch pipe passed through this flange to allow bolt clearance on the 24 inch flange. Se 24 inch flange was the mounting for the special reducing-mounting flange of the pressure vessel. Two additional 24 inch pipe sections of 1 foot and 2 feet, respectively, were also mounted on this flange for increasing the dry well volume. h e bottom p9rtion of the dry well volume was fabricated with an outside diameter of 40 inches. Be diameter of this portion was determined largely by the size of the vent pipe arrangement necessary to give adequate vent pipe area variation for tLa test program. The vent pipes were arranged to give an annular array of pipes that would simulate as nearly as possible that postulated for a prototype power plant using a concentric pool arrangement. Bere were 150 one and one-half inch diameter pipes arranged in a hex pattern with an inside clearance diameter of 16 inches and an outside diameter of 39 inches. Nipples were velded into the 1-3/4" thick " tube sheet" and fitted with couplings. Se vent pipes used were 1-1/2" and 1" in diameter. Reducing bushings were used with the 1" pipes. De vent pipes are shown in Figure IV-5. This particular arrangement shows 84 1/2 inch pipes installed to represent the area obtained with 150 - 1" pipes. IV-5

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A " water bucket" was installed inside of the dry ven to retain the water that did not flash to steam during the event and to pnvent direct impinge-ment of the water-steam mixture onto the vent tutes. It's questionable if this bucket ever retained much water. Fittings were installed at various points on the dry ven vans to obtain the pressum and temperature readings in the dry well volume. Pool and Containment Yessel his vessel, indicated as volume (3) of Figure IV-1, was a six foot diameter tank having an over-all length of 9 feet with dished heads. Se top head was fitted with a 42 inch flange that served as the support surface for the h2 inch reducing flange of the dry well vessel. De tank was fitted with numerous 1-1/2" couplings to facilitate passage for leads to thermo-couples, pressure transducers, sample tubes, power cable for interior lights, and other devices. A 3" drain was located at the bottom. A 14" pipe was installed in the middle of the tank to simulate.the pool geometry of a concentric pool. Figum IV-6 is a vecical view looking into the containment tank with the ary ven removed. Bis view shows the standpipe, p mssure pick up pipe, sample tube, and the tubing with the conductivity probe. The conductivity probe was used to give a signal when the water had been expelled from the vent tubes. Se timing trace of the Sanborn recorder saa used for recording the signal. Test Procedure The general test procedum followed a logical order of assembly, charging, steaming and firing. Se pressure vessel was " buttoned up" with the proper orifice plate and rupture disc in the 6" discharge tube. The tupture charge, consisting of an electric blasting cap and several pieces of prima-cord, was placed under the ruptun disc at this time. A thermo-l couple was placed with the blasting cap to monitor the cap temperature during the heating period. Most of the te.sts were run with water cooling l on the caps to prevent premature firing. The most successful method of rupturing the discs was obtained using Monel discs that were rated about 30% over the operating pressure with a 600 vee groove cut at about 4 5 inches diameter. Se groove depth depended upon the thickness of the disc. The firing charge was placed at a point on this groove and detonated with batteries at the proper time. l If the disc failed to tupture due to either a misfire or a weak charge the depth of the groove was such that the disc would rupture naturally at l a pressure between 1025 and 1CrT5 h e undesirable aspect of the l " spontaneous" type of operation is not knowing the time of rupture. Bis requires the chart to be run at reduced speed to assure a supply of chart paper when the event does "go". IV-6 l .-.---.,,,m ____y,,__y ,__.-___y ,-,,,,.,-,,--.._,_.,.7

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-5 The pressure vessel was filled with a given charge of water from a calibrated barrel. 21s was accomplished by pulling about 20 inches of vacutan in the pressure vessel and drawing the water in with the vacuum. Bis procedun deserated the water as 1 ; entered the pressure l vessel. An alternate procedure was to fill the pressure vessel completely with vater and drain a measuzwd amount off as the pressure built up. Be latter procedure was used primarily for the tests with samples for the entrainment studies. Heat for steaming was supplied by two calrod heating units on the outside of the pressure vessel as near the bottom (flanged end) as possible and a single unit on the bottom of the special reducing-acunting flange. Each of these units was rated at 4 7 KW - 220 YAC. Se heating period was about 4 hours, providing the internal circulation continued to 1000 psi. Every j attempt vas made to have the heater on the reducing-mounting flange turned on about one hour before the pressure vessel was charged with water. During the heating period an asbestos blanket was wrapped around the pressure vessel flange assembly to reduce heat loss at the bottom of the vessel. During the heating period, all other preparations were made. The condensing pool was filled to the proper level, cap cooling water was adjusted, thermo-couples hooked up and checked out and the pressure transducers calibrated. All of the pressure transducers were calibrated at least once a week with a dead weight gage tester. The calibrations were performed on the Sanborn recortler and test values recorded for convenient checking. Usually the recoztler was run at very low chart speed during the heating period. As the pressure approached 1000 psi the chart speed was stepped up as a-pzscautionary measure to insure a trace in the event of a premature rupture. The chart was usually run at 25 to 50 millimeters per second for normal firing. All temperatures vert recorded just prior to rupture. Be dry well and l containment temperatures were recorded just after the event. l The procedure for running the fission product entrainment samples was l essentially the same. Se tracer samples were inserted in the pressure vessel in the appropriate manner and time. Gas and water samples were withdrawn for analysis from the containment vessel and pool, h prevent I the samples of xenon and krypton from collecting in the steam dome, the samples were inserted in the pressure vessel contained in the 250 ce sample bottles with a small weep hole that would permit balance of pressure during the heating period. A plug of cerro-bend in the discharge tube of the sanqple bottle prevented loss of gas from the sample prior to the I sample being submerged, discharge tube down, in water. Se cerro-bend l melts at about 160 F to open the discharge tube to the water. When the 0 disc ruptured the pressure unbalance was sufficiently large to shatter the sample bottle and release the gas to the discharging water. IV-7 9 .---y... -.--.-%-,--em,w v,_m,2-,. -,,.,.., y -.ew .nu---w,,.,r-.,,-=------+w --+--+-weny w am.v----w ww--tm

It is intended to use this same technique with elemental iodine as the tracer. Se tests using soluble salts and insoluble particulate matter will be run with these tracers free in the water charge. Samples for analysis were collected in 250 cc two petcock sample bottles. The bottles were first evacuated and the sample drawn into them. Where possible, the bottles vert purged for a few minutes. h e gas samples were sent to the Consolidated Electrodynamics Corporation for analysis on a mass spectrometer. 4 IV-8

. =. D. Discussion of Results . The test results obtained from the transient test facility are very encouraging. In general, they indicate that the pressure response characteristics of various components of the pressure suppression system may be calculated with a reasonable degree of accuracy and conservatism. In addition, the results of tests with tracer materials indicate that the pool does serve as an effective barrier in retaining fission products. Of primary interest were the many test points run to determine the influence of system parameters on the characteristics of the pressure response in the various pressure vessels. Parametric variations were made about a given base condition. The base point conditions were those of runs number 8, 25, 26 and 27 (Figures V C-3, p V C-13, V C-14, and V C-15). Primary data for the transient tests consisted of the pressure response traces obtained from the Sanborn recorder. These traces were of transient. pressures in the pressure vessel, the dry well, the containment vessel and the differential pressure between dry well and containment vessel. Figures V C-1 through V C-20 are reproductions of these traces. The pressure and time scales have been noted on each trace. The conditions for each test may be obtained from Table IV-1. The usual chart speed was 25 or 50 millimeters (the smallest chart division) per second. There are a few at lower chart speed. These are the traces obtained when malfunction of the firing system imposed " natural" rupture of the disc. In addition to the pressure traces, temperatures were recorded just prior to and just after the incident. The temperature summary is given in Table V C-1. Inspection of the temperature data for thermocouples No. 2 and 3 will indicate the difficulty experienced in obtaining fully saturated conditions up to 1000 psi. with natural circulation. After the proper techniques were discovered, a substantial number of points were obtained with fully saturated conditions in the pressure vessel. Dry Well Pressure Resoonse Peak pressure in the dry well is the most important single design consideration in a pressure suppression system. Therefore, the comparisons of test data to calculated results for peak pressure in the dry well provide the primary basis for evaluation of the calculated results from the analytical model. These comparisons are shown on Figures IV-7, IV-8, and IV-9. One may immediately note that all of the test pressures are lower in level than the \\ IV-9

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corresponding calculated points. However, there is generally good correspondence between calculated and test values for variations in parameters. This correspondence is even better when the effect on peak pressure of subcooled water in the discharge tube has been considered. This will be discussed later. Another meaningful comparison is that for the elapsed time until peak dry well pressure is attained. Figure IV-10 is a bar graph showing this comparison using the test points for which this value could be obtained. In all cases, the test results indicate I longer time intervals than the calculated values. This is in agreement with the lower test peak pressures. An analysis was made to determine if the analytical model could properly account for the effect of. subcooled water in the discharge tube. For the test facility, it is primarily the water in the discharge tube that is released during the time interval to reach peak pressure. The energy content of this water is critical with respect to the value of peak pressure. Analytical model results were obtained using the test temperatures of runs 8 and 25 to determine energy conditions in the pressure vessel at the time of rupture. Direct comparisons could be made between test values with subcooling and analytical values with subcooling. The results of this compari-son are shown on Figure IV-ll. To accentuate the correlation, both sets of results have been plotted as functions of difference in peak 0 pressure with peak pressure at 0 F subcooling test or analytical as a base. This correlation indicates excellent agreement between test and analytical results for subcooling. It was not possible to treat all points with subcooling in this manner, therefore the graphical correlation shown on Figure IV-12 was devised to determine the correction for subcooling to apply l to a given test point. The ordinate for this figure is the peak pressure difference between the analytical model value for a given condition and the corresponding test value. As may be noted, some liberties in generalization were necessary to obtain l working data. Figure IV-13 is a comparison of the peak dry well pressures, and the time interval to reach peak, versus orifice area for corres-ponding test and analytical values. The test values have been corrected for subcooling in the manner of Figure IV-12. This curve clearly shows the increase in deviation between test and calculated values as orifice area is increased. The most plausible i explanation for this discrepancy is the reduction of flow per unit I area through the orifice as the orifice area is increased. r l IV-10 e .m,. r_..e.-,, __,_-%,..,,,e- ~ _,,,__m .. _ _ - ~ _ _. - _ _ -,, -

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e Such a reduction in flow per unit area is most probably due to steam bubble formation upstream of the orifice. Any steam in the water at the orifice will reduce the mass flow rate out of the pressure vessel. Unfortunately, the reduced area of the 6" discharge tube probably helps to promote upstream bubble formation. This situation may not be too misrepresentative of a reactor vessel

where for any break below the water line there would be baffles or other internal members in the immediate vicinity.

There are other possible effects that could cause or contribute to the deviations. One of these is choking in the shock tube downstream of the orifice due to the expansion of the released steam. This effect is easily calculated and is present to the extent that it might reduce the flow by as much as 10% for the large orifice passing theoretical flow rates. Another effect may be the condensa-tion of released steam on the dry well walls. One would expect that condensation effects, if present, to be more predominate for small orifice areas with the associated longer time intervals than for larger orifice areas. This effect is not easily calculated for the confused conditions in the dry well during the time interval to peak pressure. Uppar limit estimates for this effect corresponded to a subcooling of about 10*F. This is equivalent to about 2 psi variation in peak pressure which is not very significant. The last effect worthy of consideration is the behavior of the water as it is expelled from the vent pipes. Again, the analysis ,of the exact physical phenomenon is messy and one might expect that any discrepancies would make the test values high relative to analytical values (for peak pressure). It is possible to obtain a loose correlation (not shown)' from the test data that indicates some deviation due to depth of submergence (or behavior of water in the vent tubes). It should be pointed out that the limited number of test points does not permit extensive comparisons of parameter interactions. Returning to Figure IV-13, a simple check to confirm the relationship of peak pressures and time intervals, test relative to calculated, may be made as shown by the lines A, B and C. The value of peak pressure on the test curve may be translated to the calculated i curve to give a pseudo-value of orifice area. Using this pseudo-value of orifice area, a corresponding time i'nterval may be obtained on the calculated curve. Within the limits of test accuracy, this value agrees very well with the test value corresponding to the original orifice area. 'This check lends strength to the argument of reduced flow through the orifice and tends to confirm the accuracy of the analytical model for predicting peak pressures and j time intervals. l IV-11 i

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n.

FTGUR C TV-14

d' s 4 -vessel condenses on the walls of the dry well and creates a vacuum that is intensified as cold water sprays into the dry well through the vent tubes. There is, however, a puzzling inconsistency in that many of the tests (identical reruns in some' cases) did not exhibit any negative pressure in the dry well. One may see' the pronounced effect of the pressurized dry well on the peak dry well pressure by comparing the pressure traces of Figures V C-14 and V C-20 with Figures V C-27 and V C-17, respectively. The test peak pressures were 8 psig compared to 35 psig and 12 psig compared to 25 psig. (No correction has been made for the effect of subcooling). The purpose of the pressurized dry well is to exclude water from the vent. pipes. This, of course, l reduces the pressure build up necessary. to blow the vents clear of water. The trace of differential pressure between the dry well and contain-ment volumes indicated essentially the same information as the dry well pressure alone. This pressure gave a convenient measure of the quasi-steady-state pressure difference between the dry well and containment volumes. It is this pressure differential that forces steam to flow, during the incident, from the dry well into the pool. 2 4 It is interesting to note that the dry well pressure decay following the peak differs in characteris. tics between the runs with the large i orifice and the medium orifice. Figures V C-5 and V C-9, with the large orifice, show that the dry well pressure tends, as might be expected, to decay exponentially from the peak. The test results shown in Figures V C-19 and V C-20 do not show this effect so pronounced ' indicating that the diameter br length to diameter ratio) of the individual vent pipes may influence the after-peak pressure response. Runs 12 and 17 were with la vent pipes and rune 32 and 33 were with 1-1/2" vent pipes having an equivalent total area of the 4 1" pipes. t The question of necessary vent area is an important one for economical design. Inspection of any of the dry well pressure traces indicates that there is more vent area than necessary since l the pressure following the peak is considerably less than the peak. Run number 30 (Figure V C-17) had about one-half of the vent area (.5 vs. 9 sq. ft.) of run number 26 (Figure V C-14) with no observable change in peak pressure and only a slight increase in the after-peak pressure. i i Presumably the vent area could be minimized to the point where the pressure following the peak would be very close to the peak. For the conditions of the runs cited above, this would mean the vent area could be.1 square feet or so with no increase in dry well peak pressure. i 5 IV-13 4...

The pressure response of the containment vessel is quite interesting in that a peak pressure of 1.8 to 2 times the quasi-steady-state pressure occurs just after the drop-off in peak dry well pressure. This behavior can be explained by assuming that the air in the containment vessel is compressed adiabatically by the injection of a major portion of the dry well air into the containment vessel. And this process is followed very quickly by an evaporative cooling process due to the water thrown around within the vessel. Calculated values for such a sequence of events agree fairly well with measured values. For example, for run No. 30 (Figure V C-18) the peak was calculated at 7.2 psi versus 7 psi measured and the quasi-steady-stat

value as 4.6 psi, compared to 3.8 psi. test.

The pressure traces indicate that the air is expelled from the dry well throughout the event. The continual rise in containment pressure until the second peak is reached provides the basis for this conclusion. The measured temperature rise of the pool agreed generally with calculated values. The pool temperature rise was calculated as 0 7.5 and the measured values were 100 0 and 7 for runs 32 and 33, respectively.- A most important observation of the pressure behavior for the entire system is that all of the individual volumes return to essentially atmospheric pressure immediately following the event. This should be significant in containment leakage considerations. Fission Product Entrainment Tests The results of the noble gas fission product entrainment tests are encouraging but not particularly conclusive with respect to numerical values for effectiveness. Two tests were run with a 250 cc. (standard conditions) sample of xenon and one test was run with the same size sr.mple of krypton. The amount of xenon in a 50 MW (electrical) reactor was estimated to be about 200 gm. The test sample had about 1.2 gm. Scaling the test sample by the volume scale factor of 1000 indicates the test concentration to be 6 times that of the reactor, if all of the gas in the reactor is released! For all three tests the entire sample bottic was suspended inside of the pressure vessel in such a manner that the gas would be trapped in the sample bottle until the event was i initiated. This was accomplished with a weep hole (20 mills or so in diametor) that would accommodate slow changes in pressure. A rapid change in pressure would cause the sample bottle to explode releasing the gas into the discharging water. IV-14

l The first test using xenon (run 27) and the test using krypton-(run 26) appeared to perform in the. desired manner. The second run with xenon did not have the gas sample released as anticipated. The weep hole was unintentionally too large (approx. 25 inches diameter) and the sample bottle did not rupture. For this run, most of the gas was probably released during the last half of the event. At the conclusion of the incident, gas samples were obtained from the pool space and sent to Consolidated Electrodynamics Inc. for - analysis on the mass spectrometer. Their analyses were as follows : Mci percent Air + CO EU9El K#YDt00 2 Krypton sample 99.997 .003 + 100% - 50% Xenon sample #1 100.00 0 Xenon sample #2 100.00 0 If xenon.was present it was less than.005 moi percent which means the amount of xenon gas in the sample was of the order of their normal instrument sensitivity or less. The sensitivity for detecting krypton is better because of its unique peaking characteristics. A rough accountability of the krypton test indicates that the water pool could be supersaturated with krypton by a factor of 5 to 20. The variation is based on the accuracy limits of the analysis. The results.of the tracer tests using heavy water have not been received at the time of writing the report and nothing can be said regarding this test. E. Conclusions The results of the transient test facility have provided conclusive evidence that the pressure suppression system of reactor containment is effective in containing the energy released from a boiling water reactor incident. Furthermore, the pressure behavior of such a i system may be calculated with a reasonable degree of accuracy, and these values are conservative when compared to the test values. The primary effect that contributes to deviation between calculated ' values and test values appears to be the nature (or state) of the flow out of the pressure vessel. The analytical model assumes the flow is entirely saturated liquid. For the model tests, there is no practical method of determining the state of the fluid at the orifice during the fraction of a second before peak pressure is reached. One must conclude from analysis of the test data the degree of choking present and consider the deviation caused by choking as design margin if necessary. IV-15 L

There are three primary parameters that determine the peak dry well pressure. These are the break area, the depth of submergence of the vents, and the dry well volume. Of these, the depth of submergence is the easiest to control by design and various " tricks" may be used to reduce the effective depth of submergence. One of these tricks used in the tests was to pressurize the dry well to force the water from the vent pipes. This was extremely effective in reducing the peak dry well pressure. The gain doesn't come " free", however, as the adJitional air expelled from the dry well increases the containment volume pressure. Other characteristics of the dry well pressure response indicate several points of interest. The vent area that must be provided is directly a function of the size break that must be considered. The depth of submergence, for a given break area, is the prime determinant of what value the peak pressure will have and the vent area may be varied over a wide range with little influence on this pressure. There is, of course, a minimum value of vent area that will produce a sustained (for the duration of the event) pressure equal to the nominal peak value. However, such a vent area would produce a higher peak than a larger vent area would produce for a more serious break. The test results suggest that the vent area could be reduced by a factor of 3 or 4 with no increase in peak dry well pressure for the conditions tested. Dry well vessel designs must give due consideration to negative pressures in the dry well. For the case of a rupture in the bottom of the pressure vessel, all post accident steam generating capacity is lost by dumping the reactor water into the dry well. When the event " tails out", steam in the dry well begins to condense with an associated reduction in pressure. Cold water is sucked in through to vent pipes to promote further condensation. Measured dry well pressures went as low as -12 psi. Such behavior would prohibit the normal blowing of safety valves directig into the dry well. But such behavior could be utilized to aid in flooding the dry well after an accident. The design of the containment volume must be based on the peak i pressure anticipated in this vessel. The nature of the pressure response for this volume indicated adiabatic compression of the air in the containment volume by expulsion of air from the dry well. There was a peak pressure followed by an evaporation cooling process with the pressure dropping considerably and then a gradual rise in pressure to an intermediate value during the event. A prototype would probably not have as violent pool action as the model and less evaporative cooling. It is very significant that the entire system reverts back to essentially atmospheric pressure after the incident. This is en important consideration for fission product leakage. The fission product entrainment tests to date indicate that the pool is effective for retaining noble gases. IV-16 L

SECTION 1 APPENDIX QL21EL A Condensing Test Sample Data and Detailed Program B Condensing Test Methods of Analysis C Transient Test Data D Transient Test Methods of Analysis 't E Reference Drawings 1 i 1 5 1-n- ,--r-n...,.,.,.-e,.n..n,c,,,,,..,,..n, ._,_.wn

l 1 FIG. V-A-1 i' VERT AL II;JECTCB - ANCHORED CLOSED TA!!K COLD POOL WATER 6' DEPTH OF SUBMERGE?CE .ERTICAL SCALE-2 PSI DIV:S:0?: HORIZO::TAL 51;LE: 10 M:LLISECONDS ':!::E RUN 1

' ASS FLOW RATE 65. 3 % 4 HR.

'J12 FL3 FATE 69,700 # HR.

i ry j \\/ / i .i ( l l l I. ? s ? W PW i I L I \\. i l l l I W " ASS FLW RATE S1.700 #;HR. ESS FLOLI RATE 56.700 HR. j MASS FLOW RATE L2.300 #fHR. ESS FLOLI RATE 14.650 #, HR. l i y i f a I m.- t I I t s i ( 1 I i MASS FLW RATE 29 100 y/HR. yASS FLW RATE O #/im. i

.i FIG. V-A-2 i i 3" VERTICAL INJECTOR - ANCHORED l 1 l CLOSED TANK COLD TO HOT POOL WATER 6' DEPTH OF S'JB:GGDOE '.~ERTICAL SCALE: 2 PSI / DIVISION HORIZONTAL SOALE: 10 MILL:SE':C!CSf INCE RUN 2 0 POOL TEMPERATURE 80 F POOL TD!PERATURE 100 F 0 i s i l ,, - - ~, / '. ~.., ,_,%g l t ( { ',LO}yp x. ,. x ;, l. P 0 0 POOL TEMPERATbdE 90 F POOL TDIPERATURE 110 F 0 FOOL TEMPERATURE 115 F h ^, f s f,f (,^ l r t e t I i 1 l l l l L $'s !e 5 f. .1

FIG. V-A-3 3 7ERTICAL INJECTOR - ANCHORED 3 CLOSED TANK HOT POOL WATER d' DE?IH OF SUEMERGE';0E a yr-'!OAL SCALE: 2 PSI / DIVISION HORIZONTAL SOA'.2: .; MILLISE00I;DS/3CH RW 3 4 I

'AIZ FLOW RATE 90,000 #,HR.

!'ASE 7' N FATE 63,200 #fHR. 4 'l 'l l l N . ~ ~ ~ ~ _. _ l l l l l I i i V ~ l l l l i 1 PASS FLW RATE 79,000 #/HR. PASE 7'E RATE 55,000 p, '-3 PASS FLN RATE 41,300 #/HR.

&SS I' 3-; FATE 13,000 #,HR.

I ~ F t I n l 6 I -V v- ^'J' w ,l I i l PAIS FLOW RATE 27,600 #iHR. MAS 5 FLOW RATE O #, HR.

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i FIG. V-A-8 B" *.'OT::AL INJECTOR - UNANCHORED 0'.^,3ED TAUM HCr? POOL WATER 4 ' DEI'TH OF SUMGGE' JOE i j VERT::AL SCALE: 2 PS:,31 VISION HORIZCNTAL SOALE: 10 MILLISEJC:CS/IN:lH P L"I k j MASS FLO.? PATE 35. LOC 't /HR. l '1 1 f J I 1 ~. I U 1 l l f 4-fD 1 i i.. .l l i MASS FLOi RATE 70,300 f,HR. MASS FLOW RATE 87,200 #/HR. RUN 5 i 2' DEPTH OF SU3 MERGENCE a l f 4 . \\' 9 me \\ i t , a . c ',\\ n / ' ', y y r l l l l I r MASS FLOW RATE 72,000 #/HR. f

IABLj,y,A-1 CONDENSING IEEI, FACILITY DETAILED Il}I, PROGR AM Grouo I 4" Diameter Single Vertical Injector Depth of Run No. Flow Rate Submeroence 1 14,200lbs/hr 6 inches 2 44,500lbs/hr 1 foot 3 51,100lbs/hr 2 feet 4 53,300lbs/hr 2 feet 8" Diameter Single Vertical Injector 1 70,000 lbs/hr i foot 2 83,500lbs/hr 2 feet 3 84,000lbs/hr 2 feet 4 100,000lbs/hr 4 feet 5 100,000lbs/hr 4 feet Group II 3/4" Diameter Orifice 1 200lbs/hr Sealed Tank 2 320lbs/hr Sealed Tank 3 390 lbs/hr Sealed Tank 4 470 lbs/hr Sealed Tank V A-14

Group III a. 6" Diameter Single Vertical Injector Depth of Run No. Flow Rate Submeraenee 1 48,400lbs/hr 6 inches la 83,300lbs/hr 6 inches 2 22,500lbs/hr 6 inches Sealed Tank 3 44,000lbs/hr 6 inches Sealed Tank 4 64,200lbs/hr 6 inches Sealed Tank 5 83,300lbs/hr 6 inches Sealed Tank b. 6" Diameter Single Vertical Injector 1 35,600lbs/hr 6 inches 46,300lbs/hr 6 inches 55,000 lbs/hr 6 inches 62,200lbs/hr 6 inches 75,000lbs/hr 6 inches 2 67,900lbs/hr 6 inches 86,600lbs/hr 6 inches 3 89,900lbs/hr 6 inches c. 4" Diameter Single Vertical Injector 1 15,200lbs/hr 6 inches 2 51,500lbs/hr 6 inches 3 53,000lbs/hr 6 inches 4 58,000lbs/hr 6 inches Sealed Tank d. 14" Diameter Single Vertical Injector 1 84,400lbs/hr 6 inches 2 86,700lbs/hr 6 inches Sealed Tank e. 4" Diameter Multiple Vertical Injector 1 46,900lbs/hr 6 inches 2 90,700 lbs/hr 6 inches 3 52,800 lbs r 6 inches Sealed Tank 4 68,100 lbs r 6 feet Sealed Tank V A-15 l

Group III (Continued) f. 8" Diameter Single Vertical Injector Depth of Run No. Flow Rate Sybetroence 1 78,500lbs/hr 6 ihches 2 77,500lbs/hr 6 inches to 1 inch 3 88,400lbs/hr 6 inches Sealed Tank 4 88,400lbs/hr 6 feet Sealed Tank g. Top 4" Diameter Horizontal Injector la 24,300lbs/hr 6 inches 1 55,100lbs/hr 6 inches 2 57,400lbs/hr 6 inches Sealed Tank 3 57,400lbs/hr 6 feet Sealed Tank h. 4" Diameter Bottom Horizontal Injector 1 60,000lbs/hr 6 inches 2 61,500lbs/hr 6 inches Sealed Tank 3 63,300lbs/hr 6 feet Sealed Tank i. 6" Diameter Bottom Horizontal Injector 1 86,400lbs/hr 6 inches 2 89,800lbs/hr 6 inches Sealed Tank 3 97,500lbs/hr 6 feet Sealed Tank j. 8" Diameter Bottom Horizontal Injector 1 82,900lbs/hr 6 inches 2 82,900 lbs/hr 6 inches Sealed Tank k. 14" Diameter Bottom Horizontal Injector 1 83,000 lbs/hr 6 inches V A-16

Group IV a. 4" Diameter Single Vertical Injector Anchored to Tank Wall Depta of Run No. Flow Rate Submeroence Water 1 57,100lbs/hr 6 feet Cold (50 F) Sealed Tank 0 42,200lbs/hr ( e+at Cold (50 F) Sealed Tank 0 14,300lbs/hr 6 Met Cold (50 F) Sealed Tank 2 55,400lbs/hr 6 feet Cold to Hot (50 -120,F) Sealed Tan 6 0 0 3 54,500lbs/hr 6 feet Hot (150 F) Sealed Tank s 41,100lbs/hr 6 feet Hot (150 F) Sealed Tan) 28,700lbs/hr 6 feet Hot (150 F) Sealed Tant 14,800lbs/hr 6 feet Hot (1500F) Sealed Tani b. 4" Diameter Single Vertical Injector Unanchored 1 59,000lbs/hr 6 feet Cold Sealed Tan) 43,800lbs/hr 6 feet Cold Sealed Tani-29,200lbs/hr 6 feet Cold Sealed Tani 14,600lbs/hr 6 feet Cold Sealed Tant 2 57,500lbs/hr 6 feet Cold to Hot Sealed Tan: 3 56,300 lbs/hr 6 feet Hot Sealed Tan 43,200lbs/hr 6 feet Hot Sealed Tant 28,600lbs/hr 6 feet Hot Sealed Tan 14,300lbs/hr 6 feet Hot Sealed Tan c. 6" Diameter Single Vertical Injector Anchored to Tank Wall 1 86,700lbs/hr 6 feet Cold Sealed Tar 79,900lbs/hr 6 feet Cold Sealed Tan 68,200lbs/hr 6 feet Cold Sealed Tan 57,200lbs/hr 6 feet Cold Sealed Tac 41,800lbs/hr 6 feet Cold Sealed Tan 29,500lbs/hr 6 feet Cold Sealed Tan 14,800lbs/hr 6 feet Cold Sealed Tan 2 91,600lbs/hr 6 feet Cold to Hot Sealed Tan 9 V A-17

1 I e Group IV (Continued) d. 6" Diameter Single Vertical Injector Anchored to Tank Wall J Depth of Run No. Flow Rate Submeroence Water 1 86,100lbs/hr 6 feet Cold Sealed Tank 83,000lbs/hr 6 feet Cold Sealed Tank 71,500lbs/hr 6 feet Cold Sealed Tank / 57,200lbs/hr 6 feet Cold Sealed Tank 43,800lbs/hr 6 feet Cold Sealed Tank 29,400lbs/hr 6 feet Cold Sealed Tank 14,300lbs/hr 6 feet Cold Sealed Tank 2 %,300 lbs/hr 6 feet Cold to Hot Sealed Tank 3 90,000lbs/hr 6 feet Hot Sealed Tank 82,C0lbs/hr 6 feet Hot Sealed Tank 68,200lbs/hr 6 feet Hot Sealed Tank 55,800lbs/hr 6 feet Hot Sealed Tank 43,300lbs/hr 6 feet Hot Sealed Tank 28,600lbs/hr 6 feet Hot Sealed Tank 14,100lbs/hr 6 feet Hot Sealed Tank 4 86,500lbs/hr 4 feet Hot Sealed Tank 69,700lbs/hr 4 feet Hot Sealed Tank 5 88,000lbs/hr 2 feet Hot Sealed Tank 70,700 lbs/hr 2 feet Hot Sealed Tank e. 6" Diameter Single Vertical Injector Unanchored 1 85,800lbs/hr 6 feet Cold Sealed Tank i 4 81,700lbs/hr 6 feet Cold Sealed Tank 69,700lbs/hr 6 feet Cold Sealed Tank 56,700lbs/hr 6 feet Cold Sealed Tank i 42,300lbs/hr 6 feet Cold Sealed Tank 29,100lbs/hr 6 feet Cold Sealed Tdnk 14,600lbs/hr 6 feet Cold Scaled Tank 2 91,800lbs/hr 6 feet Cold to Hot Sealed Tank 3 90,000lbs/hr 6 feet Hot Sealed Tank 79,200lbs/hr 6 feet Hot Sealed Tank 68,200lbs/hr 6 feet Hot Sealed Tank 55,000lbs/hr 6 feet Hot Sealed Tank 41,300lbs/hr 6 feet Hot Sealed Tank 27,600lbs/hr 6 feet Hot Sealed Tank 13,800lbs/hr 6 feet Hot Sealed Tank 4 85,000lbs/hr 4 feet Hot Sealed Tank 68,200lbs/hr 4 feet Hot Sealed Tank 5 85,000 lbs r 2 feet Hot Sealed Tank 68,200 lbs r 2 feet Hot Sealed Tank. V A-18 /

Group IV (Continued) f. 8" Diameter Single Vertical Injector Anchored to Tank Wall Depth of Run No. Flow Rate Submeroence Water 1 81,900 lbs/hr 6 feet Cold Sealed Tank 82,500 lbs/hr 6 feet Cold Sealed Tank 69,800lbs/hr 6 feet Cold Sealed Tank 56,600lbs/hr 6 feet Cold Sealed Tank 43,300lbs/hr 6 feet Cold Sealed Tank 28,800 lbs/hr 6 feet Cold Sealed Tank 14,700 lbs/hr 6 feet Cold Sealed Tank 2 82,500 lbs/hr 6 feet Cold to Hot Sealed Tank 3 83,100lbs/hr 6 feet Hot Sealed Tank 81,700lbs/hr 6 feet Hot Sealed Tank 67,800 lbs/hr 6 feet Hot Scaled Tank 56,300 lbs/hr 6 feet Hot Sealed Tank 42,200lbs/hr 6 feet Hot Sealed Tank 28,700 lbs/nr 6 feet Hot Sealed Tank i 14,400lbs/hr 6 feet Hot Sealed Tank 4 85,400lbs/hr 4 feet Hot Sealed Tank 70,300 lbs/hr 4 feet Hot Sealed Tank 5 87,200lbs/hr 2 feet Hot Sealed Tank 72,000 lbs/hr 2 feet Hot Sealed Tank s g. 8" Diameter Single Vertical Injector Unanchored 1 88,800lbs/hr 6 feet Cold Sealed Tank 83,000lbs/hr 6 feet Cold Sealed Tank 70,000lbs/hr 6 feet Cold Sealed Tank 56,200 lbs/hr 6 feet Cold Sealed Tank 42,800lbs/hr 6 feet Cold Sealed Tank 28,800 lbs/hr 6 feet Cold Sealed Tank 14,500lbs/hr 6 feet Cold Sealed Tank ~ 2 88,800lbs/hr 6 feet Cold to Hot Sealed Tank 3 89,900lbs/hr 6 feet Hot Sealed Tank 82,000 lbs/hr 6 feet Hot Sealed Tank 69,800 lbs/hr 6 feet Hot Sealed Tank 56,200lbs/hr 6 feet Hot Sealed Tank 43,500lbs/hr 6 feet Hot Sealed Tank 29,000 lbs/hr 6 feet Hot Sealed Tank 14,500lbs/hr 6 feet Hot Sealed Tank 4 88,400 lbs r 4 feet Hot Sealed Tank 70,200 lbs r 4 feet Hot Sealed Tank 5 88,000lbs/hr 2 feet Hot Sealed Tank 70,200lbs/hr 2 feet Hot Sealed Tank V A-19

I.AEJ. E A,,-1, CONDENSING IIE FACILITY DETAILED IjE PROGRAM l Group I 4" Diameter Single Vertical Injector l Depth of Run No. F1ow Rate Submeroence 1 14,200lbs/hr 6 inches 2 44,500lbs/hr 1 foot 3 51,100lbs/hr 2 feet 4 53,300lbs/hr 2 feet 8" Diameter Single Vertical Injector 1 70,000 lbs/hr 1 foot 2 83,500 lbs/hr 2 feet 3 84,000lbs/hr 2 feet 4 100,000lbs/hr 4 feet 5 100,000 lbs/hr 4 feet Group II 3/4" Diameter Orifice 1 200lbs/hr Sealed Tank 2 320 lbs/hr Sealed Tank 3 390 lbs/hr Sealed Tank 4 470lbs/hr Sealed Tank I t V A-14

Group III a. 6" Diameter Single Vertical Injector Depth of Run No. Flow Rate Submeroence 1 48,400lbs/hr 6 inches la 83,300lbs/hr 6 inches 2 22,500lbs/hr 6 inches Sealed Tank 3 44,000lbs/hr 6 inches Sealed Tank 4 64,200lbs/hr '6 inches Sealed Tank 5 83,300lbs/hr 6 inches Sealed Tank b. 6" Diameter Single Vertical Injector 1 35,600lbs/hr 6 inches 46,300lbs/hr 6 inches 55,000lbs/hr 6 inches 62,200lbs/hr 6 inches 75,000lbs/hr 6 inches 2 67,900lbs/hr 6 inches l 86,600lbs/hr 6 inches 3 89,900lbs/hr 6 inches c. 4c Diameter Single Vertical Injector 1 15,200lbs/hr 6 inches 2 51,500lbs/hr 6 inches 3 53,000lbs/hr 6 inches 4 58,000lbs/hr 6 inches Sealed Tank d. 14" Diameter Single Vertical Injector 1 84,400lbs/hr 6 inches 2 86,700lbs/hr 6 inches Sealed Tank e. 4" Diameter Multiple Vertical Injector 1 46,900lbs/hr 6 inches 2 90,700 lbs/hr 6 inches 3 52,800lbs/hr 6 inches Sealed Tank 4 68,100lbs/hr 6 feet Sealed Tank V A-15

Group III (Continued) f. 8" Diameter Single Vertical Injector Depth of Run No. Flow Rate Submeroence 1 78,500lbs/hr 6 inches 2 77,500lbs/hr 6 inches to 1 inch 3 88,400lbs/hr 6 inches Sealed Tank 4 88,400lbs/hr 6 feet Sealed Tank 9 Top 4" Diameter Horizontal Injector la 24,300lbs/hr 6 inches 1 55,100lbs/hr 6 inches 2 57,400lbs/hr 6 inches Sealed Tank 3 57,400lbs/hr 6 feet Sealed Tank h. 4" Diameter Bottom Horizontal Injector 1 60,000lbs/hr 6 inches 2 61,500lbs/hr 6 inches Sealed Tank 3 63,300lbs/hr 6 feet Sealed Tank i. 6" Diameter Bottom Horizontal Injector 1 86,400lbs/hr 6 inches 2 89,800lbs/hr 6 inches Sealed Tank 3 97,500 lbs/hr 6 feet Sealed Tank j. 8" Diameter Bottom Horizontal Injector 1 82,900lbs/hr 6 inches 2 82,900lbs/hr 6 inches Sealed Tank k. 14" Diameter Bottom Horizontal Injector 1 83,000lbs/hr 6 inches V A-16

Group IV a. 4" Diameter Single Vertical Injector Anchored to Tank Wall Depth of Run No. Flow Rate Submeraence Water 1 57,100lbs/hr 6 feet Cold (50'F) Sealed Tank 42,200lbs/hr 6 feet Cold (50 F) Sealed Tank 0 0 14,300lbs/hr 6 feet Cold (50 F) Sealed Tani 2 55600lbs/hr 6 feet Cold to Hot (50 -120,F) Sealed Tank 3 54,500lbs/hr 6 feet Hot (150 F) Sealed Tan}1 0 41,100lbs/hr 6 feet Hot (150'F) Sealed Tan! 28,700lbs/hr 6 feet Hot (150"F) Sealed Tank 14,800lbs/hr 6 feet Hot (1500F) Sealed Tani b. 4" Diameter Single Vertical Injector Unanchored 1 59,000lbs/hr 6 feet Cold Sealed Tani 43,800lbs/hr 6 feet Cold Sealed Tani 29,200lbs/hr 6 feet Cold Sealed Tani 14,600lbs/hr 6 feet Cold Sealed Tani 2 57,500lbs/hr 6 feet Cold to Hot Sealed Tan: 3 56,300lbs/hr 6 feet Hot Sealed Tan? 43,200lbs/hr 6 feet Hot Sealed Tan) 28,600lbs/hr 6 feet Hot Sealed Tan 14,300lbs/hr 6 feet Hot Sealed Tan: c. 6" Diameter Single Vertical Injector Anchored to Tank Wall 1 86,700lbs/hr 6 feet Cold Sealed Tan 79,900 lbs/hr 6 feet Cold Sealed Tan 68,200lbs/hr 6 feet Cold Sealed Tan-57,200 lbs/hr 6 feet Cold Sealed Tar 41,800lbs/hr 6 feet Cold Sealed Tan 29,500lbs/hr 6 feet Cold Sealed Tan 14,800lbs/hr 6 feet Cold Sealed Tan 2 91,600lbs/hr 6 feet Cold to Hot Sealed Tan V A-17

Group IV (Continued) d. 6" Diameter Single Vertical Injector Anchored to Tank Wall Depth of Run No. Flow Rate Submeroence Water 1 86,100lbs/hr 6 feet Cold Sealed Tank 83,000lbs/hr 6 feet Cold Sealed Tank 71,500lbs/hr 6 feet . Cold Sealed Tank $7,200lbs/hr 6 feet Cold Sealed Tank 43,800lbs/hr 6 feet Cold Sealed Tank 29,400lbs/hr 6 feet Cold Sealed Tank 14,300lbs/hr 6 feet Cold Sealed Tank 2 %,300 lbs/hr 6 feet Cold to Hot Sealed Tank 3 90,000lbs/hr 6 feet Hot Sealed Tank 82,400lbs/hr 6 feet Hot Sealed Tank 68,200lbs/hr 6 feet Hot Sealed Tank 55,800lbs/hr 6 feet Hot Sealed Tank 43,300lbs/hr 6 feet Hot Sealed Tank 28,600lbs/hr 6 feet Hot Sealed Tank 14,100lbs/hr 6 feet Hot Sealed Tank 4 86,500lbs/hr 4 feet Hot Sealed Tank 69,700 lbs/hr 4 feet Hot Sealed Tank 5 88,000lbs/hr 2 feet Hot Sealed Tank 70,700lbs/hr 2 feet Hot Sealed Tank e. 6" Diameter Single Vertical Injector Unanchored 1 85,800 lbs r 6 feet Cold Sealed Tank 81,700 lbs r 6 feet Cold Sealed Tank 69,700 lbs/hr 6 feet Cold Sealed Tar.x, 56,700lbs/hr 6 feet Cold Sealed Tank 42,';00lbs/hr 6 feet Cold Sealed Tank 29,100lbs/hr 6 feet Cold Sealed Tank 14,600lbs/hr 6 feet Cold Sealed Tank 2 91,800lbs/hr 6 feet Cold to Hot Sealed Tank 3 90,000lbs/hr 6 feet Hot Sealed Tank 79,200lbs/hr 6 feet Hot Sealed Tank 68,200lbs/hr 6 feet Hot Sealed Tank $5,000lbs/hr 6 feet Hot Sealed Tank 41,300lbs/hr 6 feet Hot Sealed Tank 27,600lbs/hr 6 feet Hot Sealed Tank 13,800lbs/hr 6 feet Hot Sealed Tank 4 85,000lbs/hr 4 feet Hot Sealed Tank 68,200lbs/hr 4 feet Hot Sealed Tank 5 85,000 lbs/hr 2 feet Hot Sealed Tank 68,200lbs/hr 2 feet Hot Sealed Tank. V A-18 rw r

Group IV (Continued) f. 8" Diameter Single Vertical Injector Anchored to Tank Wall Depth of Run No. Flow Rate Submeroence Water 1 81,900lbs/hr 6 feet Cold Sealed Tank 82,500lbs/hr 6 feet Cold Sealed Tank 69,800lbs/hr 6 feet Cold Sealed Tank 56,600lbs/hr 6 feet Cold Sealed Tank 43,300lbs/hr 6 feet Cold Sealed Tank 28,800lbs/hr 6 feet Cold Sealed Tank 14,700lbs/hr 6 feet Cold Sealed Tank 2 82,500 lbs/hr 6 feet Cold to Hot Sealed Tank 3 83,100lbs/hr 6 feet Hot Sealed Tank 81,700lbs/hr 6 feet Hot Sealed Tank 67,800 lbs/hr 6 feet Hot Sealed Tank g 56,300 lbs/hr 6 feet Hot Sealed Tank 42,200lbs/hr 6 feet Hot Sealed Tank 28,700 lbs/hr 6 feet Hot Sealed Tank 14,400lbs/hr 6 feet Hot Sealed Tank 4 85,400lbs/hr 4 feet Hot Sealed Tank 70,300lbs/hr 4 feet Hot Sealed Tank 5 87,200lbs/hr 2 feet Hot Sealed Tank 72,000lbs/hr 2 feet Hot Sealed Tank g. 8" Diameter Single Vertical Injector Unanchored 1 88,800lbs/hr 6 feet Cold Sealed Tank 83,000lbs/hr 6 feet Cold Sealed Tank 70,000lbs/hr 6 feet Cold Sealed Tank 56,200lbs/hr 6 feet Cold Sealed Tank 42,800lbs/hr 6 feet Cold Sealed Tank l 28,800lbs/hr 6 feet Cold Sealed Tank l 14,500lbs/hr 6 feet Cold Sealed Tank i 2 88,800lbs/hr 6 feet Cold to Hot Sealed Tank 3 89,900lbs/hr 6 feet Hot Sealed Tank 82,000 lbs/hr 6 feet Hot Sealed Tank 69,800 lbs/hr 6 feet Hot Sealed Tank l 56,200lbs/hr 6 feet Hot Sealed Tank l-43,500lbs/hr 6 feet Hot Sealed Tank 29,000 lbs/hr 6 feet Hot Sealed Tank 14,500lbs/hr 6 feet Hot Sealed Tank 4 88,400lbs/hr 4 feet Hot Sealed Tank l 70,200lbs/hr 4 feet Hot Sealed Tank 5 88,000lbs/hr 2 feet Hot Sealed Tank 70,200lbs/hr 2 feet Hot Sealed Tank V A-19 7

Group IV (Continued) h. 4" Diameter Multiple Vertical Injector Anchered to Overhead I-Beam Depth of Run No. Flow Rate Submeroence Water 1 92,500lbs/hr 6 feet Cold Sealed Tank 84,200lbs/hr 6 feet Cold Sealed Tank 73,500lbs/hr 6 feet Cold Sealed Tank 56,100lbs/hr 6 feet Cold Sealed Tank 42,100 lbs/hr 6 feet Cold Sealed Tank 29,400lbs/hr 6 feet Cold Sealed Tank 14,400lbs/hr 6 feet Cold Sealed Tank 2 91,800lbs/hr 6 feet Cold to Hot Sealed Tank 3 92,800lbs/hr 6 feet Hot Sealed Tank 83,700lbs/hr 6 feet Hot Sealed Tank 71,300lbs/hr 6 feet Hot Sealed Tank 58,800lbs/hr 6 feet Hot Sealed Tank 44,100lbs/hr 6 feet Hot Sealed Tank 29,500lbs/hr 6 feet Hot Sealed Tank 15,300lbs/hr 6 feet Hot Sealed Tank 4 92,700 lbs/hr 4 feet Hot Sealed Tank 69,700lbs/hr 4 feet Hot Sealed Tank 5 92,000lbs/hr 2 feet Hot Sealed Tank 68,800lbs/hr 2 feet Hot Sealed Tank Group V a. 4" Diameter Multiple Vertical Injector - Tandem Compa rtment : 18" wide x 12' long x 12' deep Depth o Run No. Flow Rate Submeroence 1 19,050lbs/hr 6 feet 2 28,000lbs/hr 6 feet 3 36,400lbs/hr 6 feet 4 44,600lbs/hr 6 feet 5 51,400lbs/hr 6 feet 1 61,400lbs/hr 6 feet 2 65,500lbs/hr 6 feet 3 74,500lbs/hr 6 feet 4 85,200lbs/hr 6 feet V A-20

Group V (Continued) Comoartment: 18" wide x 10' long x 12' deep Depth of Run No. Flow Rate Submeroence 5 44,800lbs/hr 6 feet 6 52,000lbs/hr 6 feet 7 60,100lbs/hr 6 feet 8 66,800lbs/hr 6 feet 9 67,200lbs/hr 6 feet 10 72,500 lbs r 6 feet 11 79,200 lbs r 6 feet Comoartment: 18" wide x 2' long x 12' deep 12 28,200lbs/hr 6 feet 13 37,200lbs/hr 6 feet 14 46,000lbs/hr 6 feet 15 75,000lbs/hr 6 feet 16 52,000lbs/hr 6 feet Compartment: 18" wide x 8' long x 12' deep 17 40,500lbs/hr 6 feet 18 48,000lbs/hr 6 feet 19 59,800lbs/hr 6 feet 20 65,900lbs/hr 6 feet 21 73,500lbs/hr 6 feet Compartment: 18" wide x 8' long x 10' deep 22 47,900lbs/hr 6 feet 23 67,100lbs/hr 6 feet 24 81,700lbs/hr 6 feet Compartment: 18" wide x 12' long x 10' deep 25 34,700lbs/hr 6 feet 26 48,500lbs/hr 6 feet 27 70,200lbs/hr 6 feet Compartment : 18" wide x 12' long x 8' deep 28 33,700lbs/hr 6 feet 29 48,200lbs/hr 6 feet 30 68,400lbs/hr 6 feet ] V A-21 j

Group V (Continued) Compartment: 18" wide x 8' long x 8' deep Depth of Run No. Flow Rate Submeraence 31 37,000lbs/hr 6 feet 32 53,200lbs/hr 6 feet 33 72,500lbs/hr 6 feet Comoartment: 6' wide x 12' long x 12' deep 34 19,100 lbs/hr 6 feet 35 19,100lbs/hr 6 feet 36 36,800lbs/hr 6 feet 37 54,000lbs/hr 6 feet Concartment : 6" wide x 8' long x 12' deep 38 19,200lbs/hr 6 feet 39 37,200lbs/hr 6 feet 40 52,500lbs/hr 6 feet Compartment: 6" wide x 8' long x 8' deep 41 19,050lbs/hr 6 feet 42 b. 4" Diameter Multiple Vertical Injector - Side-by-Side Comoartment: 18" wide x 12' long x 12' deep 43 19,100lbs/hr 6 feet 44 36,800lbs/hr 6 feet 45 52,600lbs/hr 6 feet Compartment: 18" wide x 8' long x 12' deep 46 19,000lbs/hr 6 feet 47 36,600lbs/hr 6 feet 48 52,600lbs/hr 6 feet Compartnent : 18" wide x 12' long x 8' deep 49 19,000lbs/hr 6 feet 50 36,800lbs/hr 6 feet 51 52,500lbs/hr 6 feet V A-22

Group V (Continued) Comoartment : 18" wide x 8' long x 8' deep Depth of Run No. Flow Rate Submeroence 52 19,000lbs/hr 6 feet 53 36,800lbs/hr 6 feet 54 51,900lbs/hr 6 feet c. 4" Diameter Multiple Vertical Injector - Tandem comoartment: 6' wide x 12* long x 12' deep 55 19,500lbs/hr 6 feet 56 27,900lbs/hr 6 feet 57 36,800lbs/hr 6 feet 58 44,900lbs/hr. 6 feet 59 52,000lbs/hr 6 feet Compartment : 6" wide x 8' long x 12' deep 60 19,500lbs/hr 6 feet 61 28,200 lbs/hr 6 feet 62 36,800lbs/hr 6 feet 63 45,000lbs/hr 6 feet 64 52,500lbs/hr 6 feet Comoartment: 6" wide x 12' long x 8' deep 65 19,200lbs/hr 6 feet 66 28,300lbs/hr 6 feet 67 36,800lbs/hr 6 feet 68 45,300lbs/hr 6 feet Comoartment: 6" wide x 8' long x 8' deep 69 19,100lbs/hr 6 feet 70 27,900lbs/hr 6 feet 71 37,100lbs/hr 6 feet 72 45,300 lbs/hr 6 feet V A-23

f Group V (Continued) d. 4" Diameter Multiple Vertical Injector - Tandem Header End of Compartment to First Injector '3' Compartment: 6" wide x 12' long x 12' deep Depth of Run No. Flen Rate Submeraence 73 19,100lbs/hr 6 feet 74 36,800 lbs/hr 6 feet 75 53,200lbs/hr 6 feet 76 Check surface reaction during expulsion of air from injectors. 77 Check surface reaction during expulsion of air from injectors, e. 8" Diameter Single Vertical Injector End of Compartment to Injectors Centerline - 8" Comoartment : 12" wide x 8' long x 8' deep 78 19,100lbs/hr 4 feet 79 33,000lbs/hr 4 feet 80 53,200 lbs/hr 4 feet Comoartment: 12" wide x 12' long x 12' deep 81 19,100lbs/hr 6 feet 81a 19,100lbs/hr 6 feet 82 37,200lbs/hr 6 feet 83 54,800lbs/hr 6 feet f. 4" Diameter Single Vertical Injector End of Compartment to Injectors Centerline - 8" Comoartment: 6" wide x 6' long x 6' deep 84 3 feet 85 4,000lbs/hr 3 feet 86 11,700lbs/hr 3 ffet 87 21,000 lbs/hr 3 feet Comoartment : 6" wide x 12' long x 12' deep 88 6 feet 89 6 feet 90 10,800lbs/hr 6 feet 91 18,500lbs/hr 6 feet V A-24

V-B-1 V-B. CNDENSING TEST METHODS OF ANALYSIS Correction of the Mass Flow Rate Indicated by Barton Flow Meter: hy indicated mass rate Let = he a corrected mass rate Yg specific volume of steam upstream of the water, = under rated conditions. specific volume upstream under test conditions. T = yg expansion factor under rated conditions. = V expansion factor under test conditions. =

Then, Ed D

(1) Y [* = Wr TheratioM: V is assumed equal to unity. The rated conditions upstream are saturated steam at 100 psig. Equation (1) reduces to /, 9 7 (2) We p Wz = Velocity at Iniector Exit: From continuity considerations, the velocity in the injector at exit is (3)

s =

g, The density

m3 is unknown. Either two thermodynamic coordinates must be measured at the injector exit; or one coordinate must be measured and some assumption must be made concerning the process which the steam undergoes while flowing through the injector.

See end of Appendix V-D for definition of symbols not defined here.

1 V-B-2 For this series of tests there was no satisfactory way of measuring other than static pressure at the injector exit. Therefore an assumption was made (and subsequently, calculations were made to justify it, for the single 4 inch injector at one flow condition) concerning the process in the injector. It was assumed that the flow is isentropic but irreversible. In other words, the heat transferred through the injector wall is just equal to the energy dissipated in friction. The following calculations are made in support of the above assumption. Let F be the rate of energy dissipation due to friction and Q the rate of heat trans fer. These are given respectively by g = f l-N (4) D 29 3 and Q = A usT (5) The following are reasonable, or typical and consistent, values fo'r the single 4 inch injector, O = d, f Rs = N 780 /sec o A U = 0.4.sh &= 3 ty.sec s r = vo, l Substituting in (2) and (3), f =.5 4 #N/.sec N/sec Q=34 This is considered good enough agreement to to justify the assumption (even though at lower flows the agreement would be much poorer than indicated here). If the same friction factor and heat transfer coefficient are applied to some other size of injector the agreement is just as good. l l l l -_e-. -e

l l!-B-3 Based on the assumption of isentropic but irreversible flow, and treating the steam as a perfect gas, the following relationship holds: L .W R7~/.&i! U = p, p, Q ps/ (6) 3 This is good whether the ficw is subsonic or sonic. The only difference is that sonic flow has the possibility of being underexpanded, that is, the pressure just inside the injector at exit may be greater than the pressure in the pool at exit. Momentum Rate at Iniector Exit: Momentum rate is given simply by _N (7) Mass Flow Rate Measurement with the Sinole 4 Inch Iniector: The following is based on the assumption of isentropic but irreversible flow ( Q' = 7 ' ), and treats the steam as a perfect gas. From energy considerations ll + y. [R7; - R7~] + q ' = 0 (8) From force considerations i z 3 l 2 / *8

O (9)

Note that for isentropic flow the middle terms of the above two equations are equal, confirming that $' and r'aust be equal. l

V-B-4 The energy per unit mass F' dissipated by fricdon is /~'= UI (10) D 23 Substituting in the energy equation 3 [/ + , [}?5 -R'T;] (11) = Consider for the moment the case of sonic velocity. For this special case M3 b (12) ~- Af 2 Combining (11) and (12) y_f[ - /] (13) /* = Because the process is isentropic 1:L (14) = Combining (13) and (14) and rearranging, [/- _ 2- _d J'+/7 (15) O ~ 3* - / h

2. J

~ (Note that for reversible isentropic flow this reduces to zero) i It was found that in the case of the sinzle 4 inch injector there was a certain critical mass rate beyond which the pressure ratio S/p3 did not increase. This critical ratio is 2.1. Deducing that

fb ) coincided with the condition of sonic velocity j

this (value of at injector exit, it was substituted in (15). The corresponding value l l of f%was calculated to be 0.24 (It was subsequently found that gave better agreement with the mass rate as indicated by the Barton Flow Meter. The latter value was used in determining mass flow rate with the single 4 inch injector.)

V-B-9 Now consider the more general case again. Combining (11) and (14) and rearranging, U$ = [, y-j fl ~ ( ) (16) Substiting ap forp,- p2, and expanding the second term on the right of (16), U! = f/ f2Y f + fy*$ f)+----)(17) From continuity = A s 4 6 4'3 (18) Because the process from 1 to 3 is isentropic, u)3 = W {p W, () - f) (19) = i Expanding, "3 ~ ~ ~ ~ ~ (20) Combining (17), (18), and (20), the mass rate is given by sp}bl7-gg Abb' f(21) &=k 2M / Nole: If errors of less than 10% are permitted, then for T= 1.3 and for all AFA less than 0.17, all but the first term on the right may be omitted.

V-B-6 Calculated Mean Temperature Rise It is assumed that any change in the total mass of water in the compartment during a run is negligible. A corollary of this is that the mass rate out of the compartment is equal to the mass rate in. A statement of the first law as applied to the compartment is e i = iv[h,-E]-A U(T-7;} (22) [ is the mean temperature of the compartment, and C is where its heat capacity. Changes in have a relatively small effect on k-f, which is therefore treated as a constant. Upon integrating (22), st 4T = W4thst AU [T-T f/d (23) p It is assumed that 7is initially equal to the pool temperature T, and g increases linearly with time. s& [7-dd (24) Then Combining (23) and (24) and rearranging, k hot AT : (25) C + j gg 3 sh andl7 were taken as 1050 Btu /lb. and 0.0277 Btu /ft 0 sec F respectively for all calculations. The second term in the deonominator is a relatively small part for all compartment runs. l r

TABLE C-1 'lRANSIENT TEST FACILITY 'IEMPERATURES Time of Ruoture After Runture t raep1. Date of No. 1 2 3 4 5 6 7 Tank 5 6 7 Tank Test Point No. 6 Not recozded 1 22 7 54o 400 boo 1 8 530 45o 450 208 285 no 65 250 loo 70 1 29 9 Not m 10 Not run u 522 415 413 17o 230 78 58 185 83 72 1/31 12 534 492 490 200 253 100 73 200 72 72 1 30 13 524 370 37o 100 205 Bo 90 187 80 100 16 14 490 360 350 140 290 85 75 225 75 60 2 15 Not zun 16 537 535 540 145 85 75 75 90 80 75 2/24 17 520 M5 45 200 75 To To no 100 75 2/5 18 Not zun 19 525 43o 425 220 To 65 65 100 90 To 2/4 20 538 530 535 To 8o To To 85 85 To 2/19 21 Not run 22 Not run 23 Not run t 24 543 533 538 90 178 To To 165 91 To a 18 1 25 532 523 125 350 90 70 240 98 To 1 12 26 551 540 550 65 90 loo 2 ' 35 54o 65 375 118 loo 283 102 98 3/2 27 541 5 25 Not run 29 547 540 100 360 108 85 200 95 85 2 3o 535 Sy 518 65 2 31 500 953 492 170 348 80 75 182 100 75 2 6 i 32 542 537 542 385 n5 105 65 210 125 80 75 2/13 33 543 54o 212 330 80 100 64 190 125 100 71 3/14 nraep1. Date at Nos - 1 2 3 4 5 6 7 knk 5 6 7 Tank Test l m --en.- _., - -,.. -. - - - - - - - - - -, -.,,, n.,_

T b M. b-r.., ,.TSLawl f. .y LL. ; COL i 1 f, -N-f i b, 9 g 1; .i: ih M-_ .P. Ah; ---[ h- '* ^*3 ' "$ ~ - ['+- 'Sh 3 . i* y W... --._.., - t: } gyggg kgg,7 -w ir.a'?r.- - ~E 3* M -[. - - - - .e.w' 3.ais"w.- a g m mwm 3 2f;-@inpa = - _q g y{ ; p. i u.. g=. y =.,. e gggca. 43==.+.gtip.ggg31;gi;g;.;.g l g g; ir ..u p:;c.= g nggg=-g. y :=; magg , 3...g hi' .' y. ? NN C;

i Q. -i h

Y-hb

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x[8 w...? :

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n i ia. y.ditJ.Ei.en..t n

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ji 3- _e i 4 - F i p. 4 9 l 3 i s i 4 ( a. i r i. g . p 1 A i n t i n,,,,_ i i k. I A* i i t- h'- hi g w 1 L. -t e I i i. I I i j j -[. I Q, a a 9= h P .a.__ %___._._._ i a_.l. M M I I P I' 'I ..2._... . i. m. FIGURE V C-20

V-D-1 APPENDIX V-D Transient Pressure Analysis The maximum credible accident is the complete severance of one of the lines entering the reactor vessel or an egivalent break in the reactor vessel itself. It may be shown that the worst case is a break near the bottom of the vessel. h e case considered in this analysis is a break in the bottom of the reactor vessel. A sudden rupture of the vessel bottom results in expulsion of water through the break, which is treated here as an orifice. It is assume; chat single phase water at the state in the reactor flows through the orifice kafore it flashes to steam. Heat transfer and storage in the system boundaries is neglected in the analysis, which tends to yield a conservative result. Actually heat could be absorbed by the dry well valls which would reduce the pressure and temperature 1 inside the dry well. All friction losses are neglected. he analysis is based on one-dimensional flow. It is assumed that no mixing takes place between the pool water, the air, or the steam / water mixture. This enables the air volume to be treated as a perfect gas undergoinE reversible, isentropic changes before the pool water is completely expelled from the vent. If the air and steam / water were a homogeneous mixture throughout the dry well, the resulting pressure basedon partial pressures would be slightly greater than that predicted by the assumption of non-mixing. Calculations show that for the range of parameters we are considering, the effect of mixing on the pressure is slight. 1. "A New Approach to the design of Containment Shells for Atomic Power Plants", A. Kolflat and W. A.Chittenden,19th Annual American Power Conference, March 27-29, 1957

V-D-2 Uniform pressure is assumed to prevail throughout the dry well to the vent inlet. It is further assumed that the steam / water mixture in the dry well is homogeneous and in thermodynamic equilibrium until the end of Phase B (i.e., until the air is purged from the dry well). The pressure at the vent outlet is assumed equal to the hydrostatic pressure which exists before the onset of the transient plus the pressure acting on the surface of the pool water, i.e., the contain-ment p mssure. The containment pressure vill increase during Phase B (see below) due to the transfer of drywell air to the containment volume. De mass inflow to the system occurs in two phases: Phase 1: Occurrence of vessel zupture to expulsion of last of reactor liquid phase from vessel. (Note that break assumed in bottom of vessel) Phase 2: From end of Phase 1 to end of transient. The mass rates from the reactor and the corresponding enthalpies an obtained as functions of time. Phase 1 neglects pressure increases in the dry well. Phase 2 is based on critical velocity through the orifice for most of the interval. When the velocity drops to subsonic, the pressure in the dry well is taken to be atmospheric. Considering first Phase 1, the reactor mass rate and enthalpy are found as tice-dependent functions by a stepwise procedure. Reactor Vessel Steam _ fv Water ~ A,, Area of bretat Vu, Water Velocity ,-.---e.- e---v

The initial vessel pressure P,,

provides values for V.

and 4 k f8 Yk ' h LYv) M ' k(Yv) 4 The vessel rupture is assumed to correspeed to a sharp-edged orifice. It is further assumed that the flew through the erifice is single phase, 1.e., that ne flashing occure in the erifice2 The initial velocity of water escaping through the orifise is s

u. s

/ 2-3 v4 N ubero it is assumed that the dry well pressure P is eas1'1 in magni-tede, and that the process is quasi eteady flew for small incremsats of time. The corresponding mass rate is: M Am K .W =- 9 Successive time increments At are chosen to determine corresponding assa rates and enthalpies at the end of each interval. Eg leying the previously found W , the mess of water remaining in the reactor at ~ the end of At is, t W

We t 4t tso l

{ Ilotation appeart, lEtable at and of Appendix 0. 2. "The Flow of Saturated Water Through Throttling Orifices *g II. W. BeNamin and J. G. Millerp July,1941, ASIE Transactions.

i Employing previous vaines for both mass rate sad enthalpy, a new value of internal energy iss i E = E., + d h At 4 i s.. The specific laternsi energy of the sixture remalaing in the vessel is ) found by E es 9 W and the specifie volume by Vv Va W These two thermodynamie o*CMinates far the state of the reaotor fluid ~ at a time % = %-, + At, and the corresponding is deter-named. hn provides new values for % and' %. The precedure is repeated until V reaches the value V*3 at ubich time the iltysid phase is gone from the reacter, and the remaining steam is at 10(W quality. Considering meat Phase 2, again a stepsise processre is used to obtain l the reacter assa rate and enthalpy as time dependent functions, from the time ubes the liquid level has reached zero. It is assumed that I stasophoric pressure exists la th' dry well. Sonic velocity will occur under the condition et y f-[Ott 1 1 T-t Pv 4 Beginning with pressure Pyf I

1 e at the end of Phase 1, corresponding values for T g and

c. g are determined.

Tb

4 k Ev) 3 s = 4AM e

The ases rate for sonic velocity is given by _i h = 4A, [3.Lh b.\\ [IPI [g t 1 (V-lj (V4s ( \\ V+\\/ m The weight of steam and the internal energy remaining in the vessel are, respectively, W5 J CS W Ej AD is the difference between the immediate and preceding ~ values of P it is selected arbitrarily for each stop. AW a'ad AE are found from AP. Then At

4W E

AE ~ .W The energy rate is also given by j E 5-W l d is the specific stagnation enthalpy of the steam / water idiere flowing through the break. Solving for k, k k g AW

A new pressure increment is egleyed, and the precedure is repeated, yielding values of W and M for corresponding values for

t..

k> Sea % reaches a value such that , subsonio fiss begins, and the asse rate is given by -Q 2.3 % [ b h I ) IP ) ~ \\

Wu KAg<

i ) t 'Pn q W-0 J$ Po u ( I i The same precedure is followed as een the f1sw use sonie. The transient in the dry well takes place in three successive phases thatLAA Vessel rupture to complete expulsion of pool water from vent. M Expulsion of air from vent. thanLGA Expulsion of stesa/ water from vent. The system which we are concerned with in this analysis is composed of the steam / water, the air, and pool water which fall inside the boundary provided by the dry well enclosure, exhaust vent, and the reector vessel. An idealization of this system is shown below Vessel us11 (Rupture occurs here) Air Pool Water ~ i Dry Well Vent J Idealized Svatam I The dry well is initially filled with air, and pool water exists at j same level in the vent. t 8 + s . _ _ - -.. _... -.,.. _. ~. _,, - -

EhttLA8 A rgture occurs in the.. vessel and vessel fluid is expelled to the day well.ders it. flashes to a hemogeneous airture of steam and water. The dry sell air is compressed, and the peel ester is pushed from the AP vent. There is ao mixing between the system camponents. The air volume is bounded by the fined system boundary and the interfaces at the peel water and at the steam / water airture. Flow of the systes components l toward the exit is one dimensional. Steas/ water)g 1 I i I,Q vl Air 1 4 Ag Ail $ Pool Ester Ag p3 V. 1L m

Xst

% Vu-Vv r- -L p The asse rate of steam / water sixture into the dry well is $wa-h and the stagnation enthalpy is M, both of dich are'tiine-dependent i functions. (See Phases 1 and 2) The total internal energy transferred to the dry well in t seconds, neglecting the velocity in the dry l well is t f' J A w. s-E,s . - P v. 4t (1) i D l from the first law of thermodynaales. The total mass of water trans-ferred to the dry well is t %= v 't w d t (2) 3 ) 0 i v--- e + .__-r ---c.-,-,, _--,w-nw-,,-.,,w,- .u.-n,--,---w-------_-n,_-n,,,-a.,--- --,-w--,

V-D-8 Combinin6 (1) and (2), the specific internal energy of the water in the dry well is Ex Cy = W" (3) The quality of the water in the dry well is fcund by y" -. CN-C/ ef, (4) where Cf and efy are functions of dry well pressure. The volume of the steam / water mixture in the dry well is V8 = we n and the volume rate is found by differentiation: = WM 5p -/- fx Ng (5 y hg may be written as YN and Vs = vf

x y Wy (6)

(7) fx = Xx 79+vi,xe where is negligible. Combining (6) and (7) with (5), the volume rate becomes N)f W (Y; + Xx y) N

NW NH

+WJ H The quality rate is found by differentiating (4): "~ (9) Ny = Cff - Xy C14

V-D-v Tho sp':cific internal energy rate is cbtained from (1) and (3): c', e V ~ " " -P (10) Wu Wu The latent and saturation property ratee may be witten as bVh V45 p P (11) h (g (12) h 4 be (13) A relationship is obtained between h$ and P during Phase A by considering the air volume, whose mass is constant. WA: W ao

C o w W.

(14) The volume of the air during Phase A is t-YA dt (15) Vo Y V A* A where Y, a s - Q, (16), Assuming an adiabatic process, the air pressure is Ps P. ""\\ (17) VA / and the pressure rate is Y hr hA (18) VA t i

Combicing (4), (8), (9), (1g), (11), (12), (13), (16) and (18)', y VWaP.W. fi A v i s. W a - k+> b + ka, w" v y (, ) n- (4,xu +y ( ** Y h Va*# v, wi e Vs and P are related by further consideretten of the air end peal water. Applying Ruler's equatten to the air, E + k N + f N + fS f = o O f69+{ b')+ght+gex e (m) rat. orating.ach t.r. bet een t.e points, and r alling that f. 4(e), [p he + hM2-4d) + $(,*t-%) + b A ro (21) P

s All the air was initially at the se.e state. Assuming that it under-goes an 1sentropic, reversible process, the corresponding ther.edynamic coordinates are time functions as well as point functions. The' felle.-

ing is substituted in the fire int al: v. ~P " r S. 4. Q (,h-D d J and M! are neglected to give Beth &t P,, , d p p + D It o t f. k-t L ) 1 J O Re-arranged, ,y [oE (22) Ps: f e

V-D-11 f Impanding by tha binomial th:orea, e - v *.f. p. j ( *f +,r^g ax 1 2 re + (m) x, Just the first two terms will be used for conservatism and simplicity. DX g is next evaluated. The mass rate at any section of the vent 1. A Pw As l and A Q= f A a. Difforentiating, dM - 1O A -Sb Es fA. 6 4 f et (24) From continuity relationships, b* 4 -@ + da 4x) eda 4 a bt DA 6A Also kA As J% b 3 bt bt. uhere ks 4% is a volume element fixed in space, independent of time. Combining. ) hf A Q,1 AL bW (25) Next, combining (24) and (25), f O t en-r ,-r- - -. - -,---n -p,,. ,-+. -,py. ,---,,,,,._.m,, .m _n-,-,--,,--,,. ,,--_,.n. ,-.-,_,-n-,,

V-D-12 Assume now that g is independent of X , and substitute [A * [ which is conservative. X As z (( dX f (29) g The mass rate of the air at point 2 is 4$2 = /z #4 44 (30) Differentiating with respect to time,

s S

u b #

dtgg, 4'zh (31) di dl #' #'
d 4

or d%z = gg- $tV +Vzhg 2 (32) i

Comb..ning cith (29), At As f, d '

g g 1 il

= g 4x vs +Vs n r, K.s A t Ks A rL ks t hs d dA

4 4

g (33) ds p At k Substituting k = fo ,(33)becomes gg. ' h*g is 4 s. Ps. 4. A* (34) L Y x, 4 We here assume As

1. a itnear functicaaf Xs As= A. - x A 'A t (30 from which At Os Mk As (36)

A j Ai Combining (34) and (36) with (23), aestning that E = E-9%

2.

k '+ g An expression relating Ps and P3 may be found by applying Euler's equation to the pool water X3 w, N.4 + S - A + y,t (4-m') +y [$ D ^ =a(30 Ws The mass rate of pool mater at point 2 is Wh* f N 1 I free uhtch h asas (39) As

v u-14 Differentiating cith respect to time, recalling that As is a function of is only,andcombiningwith(35), e. DsV h. 1 (m) h As A.- 4 %s Prem sentinuity, i uAs3s h As b W Mg s

s. -

%[- Wd * ~I 3 but A=h* $s - us At ] t 5 As, Therefore, ,s "I (41) M3-4s 3 combining (41) and ( e) with (3s), .s Ps-Tb 'WP (h'*s) +[-I hk -

('

(42) 511mipattag Pg by the combination of (37) and (42), the resulting expression una selved for s and integrated to obtain an expres-sien.f,or hs. t -Xs was substituted for Es-1,. f., t UA~ '5 ~ P g ~, p L -35.9 $ -i A(0 -A (^i,) ~ ~ + ebp a.. q 1 my') = v' dt (43) (h.f Y, ~l M(hM k 3.. Phase A is ended uben )(s.L.. = v-

thaan 1: Pool water is gone from the vent and air starts to be ejected into the pool. E .' Steak /.. I i 4 t.r, 1 Air A v 3 3 / \\ P,' Pg' , measured at the throat of the vent, is greater than or equai to the critical pressure with respect to P If P is great

enough,

?' la critical and 3 s f 2. E-t p3 4 p3 -1 (44) y e W+\\ Considering the subsonic case, the mass rate of air in the dry well is fw A3 u s f3 - (e) Re-arranging and substituting f1, s I f (9 s Q3 (46) As LA s is given by ~g u,-Rfh-T I~ i)4 g gF L (} where A is the sonic velocity under stagnation conditions. /t 9v g (48)

o. =

/Y 9.T = O O a .,m ,-----------.--,,-,,n--_-,n-,, ,,_,---e,,n,--.-----,-s-.,,---------,n----e e--r+> m--

Assuming that the air endergess a reversible, isentropic process, /Ed m v_- \\D. v and solving for P 'If, Pv" = P. V. 9 (49) Substituting (49)into(48), (M) l 4= 3 3 P. "If, P Combining with (47), u,.lw-i 3e.b.'e B,. i- (9)y .1 (51) J W Substitut1'nginto(46), 4 l-M (52)

i h Y.

P = In order to consider the sonic case, the fellowing substitution must be made for in(52): y, tt v-t 9 Q 4\\ which gives 4 1Lt) y a-i su h. Yg?.g 23 i P (m) v. A3 +L t j Iqpations (52) or (53) are solved for h. Equations (8), (9), (10), (11), (12), and (13) are combined and arranged to give an expression 1 for P eich is then integrated to give P. The effect of a steagwater mixture entering the vent will be neglected m_._,,- ,,y,,,. ,-.,_m. _.__,.7 -.--,-__.__--_,-.,-_-.__,v_,y_- - - -,_--r-._

V-D-17 until all air is expelled, or until J t VN fMd (5h) = at which time Phase B is eJed. Coincident with the expulsion of air from the dry well is the addition of air to the containment volume and the attendant rise in containment pressure. 'Ihis produces a corresponding rise in the pressure at the vent exit. [c fMA (55) [y = p The containment pressure is increased slightly during Phase A fc t

h* (

) (56) VC refers to the volume of air above the pool, which is reduced during Phase A due to the expulsion of water from the vent. Let T be the fraction of dry well air at any instant which has mixed with the containment air. Assume complete adiabatic mixing. Then, per the first law, Ec = Eci + r Wso Asc }'_j'[((,") + rWas 4, (57) = Eco + But Ec = frWoo t We, Cc (58) //e,= P. W

d '

(59)

V-D-18 Combining (57), (58), and (59), Wco{Cc ~ Cco ~ f'*'f(hc') i ) = rWso[R&-(ee -c'.)](60) y-It is assumed that the initial containment pressure and temperature are the same as for the dry well. Sen de. C,,A,*ja0 'Fo Letting s$ e = c a[ and substituting in (60), '2r, and = e r VeofCy(5-TA) y_ ) ff = r Vpo& -c,(7, -7, A This may be rearranged to give f* Vc, t V'* Ve, + Von ~ 1 l vhich reduces to i { Tc r W. g j #~' ve, y Nc* tt Voe, \\ Vei / %, + r no (61) y he containment pressure is given by b - (YW* +M*])?TE _ [' 7, "Tc~ % (r%o+ %o) j Vei W, (62) ~ j combining (61)ana(62) l A = ^ v? [t 'C.' + G~)~l p.[r %' + (t)'] () = At the end of Phase B the containment pressure is given simply by fc = fo[Y , ' + V',*)] (64) The pressure at vent exit is given by equation (55). i i l

V-D-19 Phase C: The last of the air passes out of the vent, and steam starts to be ejected. This part of the transient is treated as quasi-steady state. Two special assumptions are made: (1) he steam / water is no lon6er homogeneous, but undergoes complete separation in the dry well, and (2) only the vapor phase flows out of the dry well. S e pressure is uniform throughout the dry well at any instant, and flow through the vent is reversible and isentropic, both assumptions made pre-viously. Let the rate of heat transfer from the dry well be Q. Ren, treating the dry well as a themodynamic systes and applying the first law, 8g Ny/r - M f -- Q (65) y but En =WC +We (66) y y 4 p (Note that o is the saturation value corresponding to f the dry well pressure p. But the aversge internal energy of the liquid phase in the dry well may be different. In this analysis the average may be higher or lower than the saturation value. This assumes negligible flashing in the event of a drop in dry well pressure.) he total steam / water volume /g is equal to the dry well volume, hence constant. Therefore a V ~$ H f $/ g/, y,, or 1 (67) W=- j

V-D-20 From conservation of mass principles g W, 4 W s = l,l - W, (68) W -W = u 3 Combining (65), (66), (67), and (68) el j< e' h 8 ~ W. W N *Nh -Q = y (69) u sy s' hy/) - Q t N (70) Let Vs e' - U] e1_ f ) (71) /-% 3 l1, ~ $ {f ~ 3 f s Then X - 6[p) : b(# (73) [ /g and b are given functions of time. Q can be estimated. M and [ y are known, hence h is a function y y of [, permitting the representation, as given by equation (72).. Equation (73) must now be solves for p. A convenient way of doing this is to plot p-fj, versus Nu for a given value of [7 and A , and various values of _ Y. y In preparing the plot,hy must be detemined. From energy considera-

tions, p ' f2 (1,,- h,)

(74) W, = f State 3 is at the vent exit. It is determined from the pressure and the entropy 5,, given by i 5 5 (7s) ls, ~ Sg, 3 N ~ 3 - 5,9 y (76) 1 7 J3 (77) v',

Y, Uij s 4 "is

l V-D-21 Results The equations for Phases A and B have been arranged for numerical solution by an IBM 650 con:puter, and for Phase C by desk calculator. Solutions have been obtained for the same parameter combinations as employed in the Transient ' Inst Facility. Comparison between these and the test results appear in the text of this report. L

V-D-22 NOTATION Dot notation is used throughout to represent the time derivetives of functions. Symbols M Area (sq. ft. ) d Sonic velocity, stagnation conditions (ft/sec) f Totalinternalenergy(ft,lb.) C Specific international ener6y (ft. lb/lb) [ Function of pressure G Mass flow rate per unit area (lb/sec/sq. ft.) f Gravitational constant (ft/sq. sec) M Enthalpy (ft. lb/lb) Ne vent constant (ft) 2 Submerged length of vent (ft) 8 Pressure (lb/sq. ft.), dry well pressurewithout subscript Q Heat (BW) 8 Entropy (BW/deg.R) T Absolute temperatures (deg.R) 8 Time (sec) 4 Velocity (ft/sec) V Volume (cu. ft) V Specificvolume(cuft/lb) W Weight (lb) Al Weight density (1b/cu. ft.) )(g Quality of steam / water mixture in dry well (vt. fraction) Xf Pool vater displacement from original level (ft) E Pool water elevation above arbitrary datum (ft) T Ratio of specific heats

~ V-D-23 Symbols (continued) f [ Massdensity(Slugs /cu.ft.) 3 l l Subscripts / Air t 8 Break in vessel C containment vessel j f Liquid phase [y Vaporization f Vapor Phase I // Dry Well (absence of subscript may also refer to drywell) i O original state before onset of transient l [ Pool j V Reactor vessel f 1 Inlet to, or upstream of, vent, end of Phase A f ] 2 Interface between pool water and air, end of Phase B [ l j 5 Exit of vent 4 i i ) l l I l j i i a i .j.

~~ ad O y r. e fi i

s o

l 2, t e.9 n.e gt.4f,u 3 C 5 i, 5 P ,i o, o ll _$ l al,$&.4.).{[ .r'." c .l

r)I irea

/t g :. ~: Ao M

i."

c F@ N i

trer: is i

l mi t N: : I ,t. a -l h i ~ J l 3 'i' l-P ~{ uf r., 2 50 j! LI)r j1n P ,l L zgu te i l I i: U, I I i l' W1 r I I . 'l l* i 1 i e -)l i i 4 i i a, n.....J w s b i

j 6

i l I l i l i J h l ', s* i l h ) e I.k 8 1 i + 's = j 'I.t t. l i i I .-6 l ~~'a g m.o- [r,.ty r> 1 ; i i i I 1 I 9: r t / } l i $? h.. g Q' r,T .] I !. r-t_ r p i g, _._\\\\. "'y]9 u,. { ( l @--~4 f.\\. l c . l -, L+ 1 a h. } CC.

4.

I 1 'l 15 .a e Y .s 59 e ,u 3 e 5 $t 11 '1 Q$ + n &A 1

L**"**_

h s 1 s u (7i ! d is O si o 3 L.f.l.f. 3 n.* 434 O ?(I T7 ~ l \\ V, 4 ii, n,E ai e 1 ,g a v .= =, gs ujo w$ t 0[*.v..n y J'k l 'I,? N)@ 7 1 .9 ~ g g v 4 utM 1. i c o: -., e a s. v 24 t 1 l x .g3e l u l + g6 F ,,i r d dt hj 0 g l ) I I J ii__ a-i I I I (s.'.-- t a. l l ] 5 l l d .,1 1 a - r.-.s... _ _.,_. s i 1 l 3 i I l l f r l 1 n i .) l-J 1 j i o ;1 g 3 3 i ~~\\ l q, a ->- t l l 4-r> u 1 vt I lE d m ' y l g l

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tr g.'a k

.T g n 't. 9. ; r L _ _,..... m c i q M\\ ) vn e k0

5 h

Pi I b $t it p 5 'f s e, 44 I s e = - -..._ -__. ,_, __ _ L = - __ --~v-mmw 9

i l I l t*! d il o l D F i 47.1 i f=ll o

'y$g5 TITI,J-.

'./ '3 1 g ai!g:!'s

1m :: li st o

E d I A 84v l :... : : we

g: 4'

'c e b n l f i 1, a c. r I Y 1 J EI >![ ~ za

p

( j r-i l rt-x \\ a l'.is0 E l E l' 3 ~ 56 ji.~ ll' ll,L I i lii i i r___ I i l ), !l 1 Is l-~~ ll' ,l l

.,.;; - l, ;-

i i i -..,.. s l l 7 ii 3 i / i i ~ I l .l ! l t 1 l 1 i I a ki e 4 s a-3.n>. ~' r' 1[" A f) -M-a, ~ I 'V 'i -qn,7 i p O q; ^ l r q_www h_<, n i eW t* W = l 'k .,T.i t, cm lS ~ i 1 \\ p\\ L .) Mp i t' e as TS s' $ >t I U i ,11 .-, g A s$ 6 I

__%m a---we __m.-_._ i i l i I I 4 - _.' ML E d l, j.++++1 l a 'e!. O y' I ( l o I .N, t. i I L_ }, 3 e / w{ ,1,,, l 7 %.,+,, dim.t + ; \\ !:,M .r 4 E, I v - =w {. g '., /~ 3, i .g-. - - v.. i l -. 3 3,g - ._ J l I ~ j g I l p h A k'g-(~,g i o i {/ [.. I / L ,,_ l,. 1 f ) i a A l ig. s...sc.C. I 1 / i ,),,;,_2,'g)E.,./. - I.t y ~ 4 i ,,_ q ef*Ie.L);- M8 -e .c..c .f.] p,y..f[U f .[;,,,, ;-... CIj ..o.3f.. s.w -- o ..m.c . i n I . ;.y 3 ,e P 6 es w .7 w l. l ,g !,./'. s.J f J ',,.,/ y I I 4 /, - Tjh3I ! N.w!hd. 3 i _wl- -2.a q.,, .a== t': .6 j -f kI f!f i l ' " ' * * * *. "i 1i 1 Ly onj .g;.6, j

7 +'g - -

li ((,-)) i _.' ~ + t 3 t.,0 *.3 -~~~-.,. 4. p h JM., e < ';s i t

N,,'.,[

I, I

[,IM,/ Q I

a g,pcqi { m- .._...y. s.;, .o w l o.. t, A.... t, 3 > . n ....-sy,,.. i 'r t V , j,, g ,.u4-n rW,tt ge Ifi- -.,,,--r, - ; r, r,- ,-,f,--, l

CS cp j}d;iXi-$ L i 'e (it 4

f..

i 3 r .L,.. _..p { / l gl t t.. n,1, i.,. 7I weu., + ). .f < .-8:eE J.4C 1 3 j .n l 1 4.. J ./ di(

a.,,

p.1; -?~ es l _. 3 3,s. s. / s s*-(>,['[)i l h4 -.\\ / M 1, !.L

_.}

y Q li .i l y .mua.,_ R /c,.,. -(.k vy g)1./}.1

i

,h t 3 '

y_[,{-h 1 L - ](l[ji' a.m z,. +. i,0 ' I ~. n,4 --~ v n

..... h, t! 9. ...';;if i ' .. c F,l f,. I :.: ' t l J rv ' "f p',*' . ', 1!! u

@. " t.'.:. U j.r M[1

'[ ,-.2 l' I li D i,,

nf. +n.:.n. e.,.4 m.

> -w x -.. s ..,t,s,,

- - [f f -- -4{ { ~d W*-

erH -i c.g. /'! pl 1 9,.,,.d.., :,y o_u. 4 - ,1 (. r_ - s\\ .-; g g....w - ~.. Q\\ ..,.. J.f s wo ...ef.efgt, i f T:/,:?,'Q{Qj[t < ty <%g t ...., ne..; u ,unt ~ u- ~ =.:: I.t,,-m,.,,w.r.i 4 f N I Ib, w n, .I h.r c rm ;r. m + '. T i r t- . T --'. i . i 1

es,.a me_2 =am.m.w. m.--

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Dear Mr. Menz:

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