ML17334A364
| ML17334A364 | |
| Person / Time | |
|---|---|
| Site: | Cook  |
| Issue date: | 02/17/1982 |
| From: | Hunter R INDIANA MICHIGAN POWER CO. (FORMERLY INDIANA & MICHIG |
| To: | Harold Denton Office of Nuclear Reactor Regulation |
| Shared Package | |
| ML17319B196 | List: |
| References | |
| AEP:NRC-00500G, AEP:NRC-500G, NUDOCS 8202240166 | |
| Download: ML17334A364 (121) | |
Text
REGULATORY ORMATION DISTRIBUTION SY M (RIDS)
ACCESSION NBR:8202240166 DOC ~ DATE: 82/02/17 NOTARI'ZED: NODOCKET, FACIL:50"315 Donald C.. Cook Nuclear Power Plantt Unit 1P Indiana L 05000315 50 316 Donald C. Cook Nuclear Power Plant~ Unit 2P Indiana 8 05000316 AUTH ~ NAME AUTHOR AFFILIATION HUNTERER.S, Indiana 8 MichiganElectric Co.
REC IP NAME
~ RECIPIENT AF F IL'I ATION DENTONEH ~ RE Office of Nuclear Reactor Regulationi Director V V ~
SUBJECT:
For wards addi info re hydrogen m$ tigatjon 8 contr ol
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Responses to SA Varga 810715 8, 0904 )'tr sir esultts of Amer'ican Electric Power CoEIncg'TVA 8 Duke, Power <<Co experimental Itest programs 8 proposed 'Tech Spec CODE: A001S, COPIES iEECEIVED:LTR, encl'ISTRIBUTION Distribution for after -Issuance of Operating License
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INOIANA II MICHIGAN ELECTRIC COMPANY P. O. 8OX IS 8OWLING GREEN STATION N EW YORK, N. Y. 10004 February 17, 1982 AEP:NRC:00500G Donald C. Cook Nuclear Plant Unit Nos. 1 and 2 Docket Nos. 50-315 and 50-316 License Nos. DPR-58 and DPR-74 HYDROGEN MITIGATION AND CONTROL PROGRAM Mr. Harold R. Denton, Director Office of Nuclear Reactor Regulation U. S. Nuclear Regulatory Commission Washington, D. C. 20555
Dear Mr. Denton:
This letter and its attachments provide additional information regarding Hydrogen Mitigation and Control. at the Donald C. Cook Nuclear Plant. Attachment Nos. 1 and. 2 provide additional responses to Mr. S. A. Varga's letters of July 15, 1981 and September 4, 1981, respec-tively, and supplement the information contained in our AEP:NRC:00500F-submittal dated October 28, 1981. The results of experimental test pro-grams co-funded by American Electric Power, Tennessee Valley Authority and Duke Power Company are included in Attachment Nos. 3 through 6.
Attachment No. 7 contains the results of Cook specific core recovery studies. Attachment No. 8 contains revised Technical Specifications for the Distributed Ignition System. The enclosure to this letter contains a listing of the material transmitted herein.
'his document has been prepared following Corporate procedures which incorporate a reasonable set of controls to insure its accuracy and completeness prior to signature by the undersigned.
Very truly yours, R. S. unter Vice President
/md cc: John E. Dolan Columbus R. W. Jurgensen W. G. Smith, Jr. Bridgman R. C. Callen G. Charnoff Joe Williams, Jr.
NRC Resident Inspector at Cook Plant Bridgman
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En'closu'ret'o'EP:NRC 00500G Lis't'i'n 'ofttachme'nt's
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Additional Xnformation Responding to Mr. S. A. Varga's letter of 15 July 1981.
Additional Xnformation Responding to Mr. S. A. Varga's letter of 04 September 1981.
Report on Hydrogen Combustion Control Studies performed by Acurex Corporation for AEP, Duke, TVA, and EPRX.
Report on Xgnitor Performance Studies performed by Whiteshell Nuclear Research Establishment for AEP, Duke, TVA, and EPRX.
Report on Hydrogen Combustion Phenomena Studies performed by Nhiteshell Nuclear Research Establishment for AEP, Duke, TVA, EPRX, Ontario Hydro and Atomic Energy of Canada Limited.
Report on Hydrogen Mixing and Distribute,on Studies performed by Hanford Engineering Devel'opulent Laboratory for AEP, Duke, TVA, and EPRX.
Report o'n Cook-specific Core 'Recovery Studies.
Revised DXS Technical Specification.
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8202240166 Attachment No. 1 to AEP:NRC:005006 Donald C. Cook Nuclear Plant Unit Nos. 1 and 2 itional Information on Hydrogen Mitigation and Control Additio Responses to Mr. S. A. Varga's letter of 15 July 1981
Item 1 Attachment No. 8 of this submittal contains revised Table 3.6-1A for the Distributed Ignition System (DIS} Technical Specifications.
The attached tables reflect the as-installed configuration of the DIS for Unit Nos. 1 and 2 of the Cook Plant and supercede the tables transmitted via our AEP:NRC:00500C letter dated 29 May 1981.
(Reference 1) .
Item 3 The Tennessee Valley Authority (TVA) has conducted tests at their Singleton Laboratories to determine the effects of lean (0-4 v/o) hydrogen mixtures for sustained duration (24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />) on igniter performance. The results of the TVA tests are contained in Reference
- 2. We have reviewed the test results and concur with TVA's conclusions.
That is, prolonged operation in a lean hydrogen mixture did not inhibit proper functioning of the igniter.
Item em 6 A topical report on the CLASIX computer code (Reference 4) has been prepared by Offshore Power Systems (OPS) and previously trans-mitted to the NRC by TVA (Reference 2) and is not being retrans-mitted with this submittal.
a) The requested information is contained in sections II through VZ of Reference 4.
b) (i) The requested information is contained in Attachment E of Reference 4.
The requested information is contained in Attachment F of Reference 4.
(iii) The requested information is contained in ~
Section VIII and Appendices A, B, and C of Reference 4.
(c) The requested information is contained in Sections III, IV, and V.D of Reference 4.
(d) The requested information is contained in Sections II, IV, V.A.l, V.D, and VI of Reference 4. The CLASIX input values used for Cook Plant analyses are given in Table 8 of Attachment No. 2 to our AEP:NRC:00500E submittal dated 02 July 1981. (Reference 5).
(e) The requested information is contained in Sections IV, V, and VI of Reference 4.
(f) The -version of CLASIX utilized for the analyses contained in Reference 6 does not rely on the LOTIC 1 computer code for initialization as the code (CLASIX) runs from the onset of the transient. The initial containment conditions assumed in the CLASIX analyses are consistent with those used in the FSAR containment analyses.
(g}. The requested information is contained in Sections II, IV, V.C and VI of Reference 4.
(h) The potential for preferential flow to the ice bed (mal-distribution) following an accident was investigated during the design of the Cook Plant and found not to be a problem.
Numerous analyses were performed to address maldistribution due to small breaks, nonuniform ice melt and jet impingement on the lower inlet doors. The results of these analyses indicate that 'burn through'as not a problem and that considerable margin remained in the ice bed. The results of these analyses and the phenomena associated with potential maldistributions would be comparable for the hydrogen-air-steam mixtures presently under consideration as for the steam-air mixtures originally evaluated and found to be acceptable.
The arrangement of the reactor coolant system precludes the hypothetical break being adjacent to the lower inlet doors of the ice condenser and should result in substantive
/'ixing in the lower compartment prior to entry into the ice condenser. The research program at the Hanford Engineering
Development Laboratory (HEDL) was designed to evaluate the extent of mixing within the lower volume following a small break LOCA. The results of the HEDL tests should provide additional evidence of mixing within the lower .
compartment and should also verify uniformity of the lower compartment atmosphere following a S2D event.
(i) The requested information is contained in the cover letter to Reference 4.
Items 7 and 8 The requested information and analyses will be transmitted at a later date.
Items 9 and 10 A study was conducted to identify the major fog formation and removal mechanisms within an ice condenser containment. Additionally, the effect of any resulting fog on the operation of the distributed ignition system was evaluated. Fog concentrations throughout containment were obtained by simultaneously solving the mass con-servation equations for fog droplets in each of the containment subcompartments. These equations incorporated fog formation due to condensation and reactor coolant system blowdown as well as fog removal by gravitational settling and sprays. The conclusions were:
- 1) Calculated Fog concentrations based on Clasix work will not affect the hydrogen flamability limits in the lower and upper com-partments. 2) Although unlikely, the expected fog concentration present in the upper plenum may raise the lower flamability limit slightly.
Analytical work has shown that the mean droplet size formed by the break flow is approximately 4p. To account for condensation and agglomeration, this study assumed a mean droplet diameter of 10p. Further analysis indicated that significant, amounts of fog are generated only when the initial state of the break flow is subcooled or saturated liquid. Therefore, fog generation was
assumed to terminate when the reactor vessel level fell below the break elevation. Approximately fifty percent of the break flow was vaporized, leaving the remaining mass suspended as fog droplets.
Knowing that thermal boundary layers are formed ne'ar cold surfaces, the Hijikata-Mori theory of fog formation within the thermal boundary layer was used to evaluate fog formation due to nucleation.
Analysis has shown that when local vapor supersaturation reaches critical supersaturation, rapid fog formation occurs within the boundary layer. Since micronsize droplets are obtained within a few milliseconds., very little supersaturation is required for further growth. During fog formation within the boundary layer, the bulk fluid temperature is decreased via heat transfer to the cold surface.
If the bulk fluid temperature has dropped below the dew point, then it was assumed that bulk stream condensation has occurred. Calcu-lations based on studies of the growth of cloud drops by condensation and observations of valley fogs have shown that the mean droplet size varies from 8p to 14'. A 10p droplet diameter was, therefore, assumed.
The only significant fog removal mechanism considered was collision with spray droplets. A removal efficiency of 100$ within the volume covered by containment sprays was assumed. (Varying the fog removal efficiency did not significantly affect the study's conclusions). Although removal by gravitational settling was elevated, the terminal velocity of a 10p droplet is so small that the effect was minimal. Agglomeration is a potentially significant removal mechanism that was simplistically treated by assuming the mean droplet size grew to 10p. Other potentially significant, re-moval mechanisms that were not considered are impingement on structures and the "spray" resulting from ice melting.
The effect of a fog on hydrogen combustion was evaluated by utilizing the work of von Karman. Three energy equations were
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developed to model'he'ombustion kinetics and heat transferred to the unburned gas and fog. Comparisons between von Karman's work
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and experimental. data obtained from Factory Mutual were good.
Analysis has shown that the minimum fog incr'ting concentration varies approximately with the square of the volume mean droplet size.
Analysis of fog formation and removal rates revealed that the upper and lower compartments maintained lower fog concentrations than the upper plenum. When the hydrogen concentration reached the lower flamability limit (4%) in the lower and upper compartments, the calculated fog concentrations were a factor of ten and five, respectively, below the inerting limit of approximately 8 x 10 v/o. The calculated fog concentration in the upper plenum was sufficient to inert the mixture when a hydrogen concentration of 4 v/o was achieved; However, when the hydrogen concentration reached 8.0 8.'5 v/o, the calculated fog concentration was at least two times smaller than the required concentration for inerting, approximately 2 x 10 v/o.
Because of the conservatisms mentioned earlier, essentially all fog generated within the lower compartment was transported to the upper plenum. The severity of these conservatisms was obvious when the analytical model was compared to the complex geometry and congestion present within an ice condenser containment. However, due to the complexity of the analysis, fog removal by mechanisms other than spray collision and gravitational settling was not quantifiable. As a result, it was recognized at the beginning of the study that the calculated results would be more qualitative than quantitative. Considering the calculated results and the conservatisms inherent in the analytical methodology discussed earlier, it was concluded that: 1) Pogs present. in the lower and upper compartments would not affect the hydrogen lower flamability limit. 2) Fogs present in the upper plenum are not expected to affect hydrogen's lower flamability limit.
Item ll (a) The following response, was provided by Dr. Bernard Lewis and Mr. Bel'a Karlovitz:
"As soon as the mixture becomes flammable it will be ignited by glow plugs in the upper plenum and the flame will move downward into the ice condenser along the boundary layers.
The flame settles to a level where the mixture turns flammable owing to water vapor condensation. Xf the mixture is already flammable in the lower compartment, it is ignited there by the glow plugs and only burned mixture passes through the ice condenser. The ice condenser is never filled with a strong H2-air mixture where the H2 concentration approaches the detonable limit.
Measurements of detonation velocities in mixtures with various H2 concentrations show that the detonation limits are sharply defined. At the limits, the detonation velocity drops off sharply "to sound velocity, creating a weak pressure wave . lt follows that outside of the detonable limits no transition to detonation is possible for any geometry of the system. To prove this statement, experiments at various geometries should always be carried with mixtures within and outside of the detonable range.
Results obtained for transition to detonation under a given geometry of a system are applicable only to similar geometries; for example, values of L/D obtained in a smooth-bore tube are not applicable to a rough surface tube or to more complex systems.
The complex geometry of the ice condenser could not be simulated by a simplified test system".
See, for example, Lewis and von Elbe "Combustion Flames and Explosions of Gases," 1961, Figure 279, page 530, and Figure 284, page 536.
(b) The 'following response 'was proVided by Lewis and Karlovitz:
"The flame will spread in ever'y direction 'from the glow plugs in the upper plenum. As the flame reaches the'op of the ice condenser baskets, it, will move. downward in the boundary layer of the stream. Structural elements crossing the stream will act as flame holders and carry the flame across the stream.
Ultimately, flames will settle in a zone where sufficient water vapor is removed to render the mixture flammable.
Flashback of flames 'into tubes was one of the earliest phenomena which has been studied extensively. The experimental results are summarized by recognition of the fact that flashback is controlled by the velocity gradient of the flow in the boundary layer at the wall 2 . The velocity gradient of the flow in the boundary layer can be calculated from a diagram which was originally published by Prandtl and Tietjens and. reprinted in the Princeton Series on High Speed Aerodynamics.
Assuming equivalent diameter of 50 cm for the channels between the ice baskets and one foot (30 cm) average flow velocity, the Reynolds number of this flow is 11,000, and the velocity gradient in the laminar boundary layer is 42 sec.
This velocity gradient is to be compared with the critical boundary velocity gradient for flashback given on Figures 93, 99, and 100 in reference 2 which show that for flames having a laminar burning velocity of about 40-50 cm/sec, the critical boundary velocity'radient is in the order of 400 to 600 sec. -1 Accordingly, flames of even very weak hydrogen-air mixtures will flash back in the boundary layer of the flow in the ice condenser".
2 Lewis and von Elbe ibid 1961 edition, pages 243-253.
Prandtl and Tietjens "Applied Hydro-Aeromechanics," McGraw-Hill High Speed Aerodynamics, Vol. lX, "Combustion Processes" Princeton University Press, 1956, page 359.
(c) The EPRZ Whiteshell test plan, which was established in early March 1981 after extensive discussions among EPBI, AEP-Duke-TVA, Whiteshell-AECL and Ontario Hydro, reflects the several objectives that the group set out, to realize. The details of the Whiteshell program have been related to the NRC through quarterly research reports and discussed with the staff in person on different. occasions. Here, the key elements of the research program which are pertinent to the issue of hydrogen combustion are summarized.
The principa3. thrust of'he Whiteshell effort is focused on providing confirmatory information in support of our selection and implementation of a distributed ignition system and in augmenting the current understanding of hydrogen combustion phenomena which are pertinent to the application of that system. We have maintained, based on the recommendations of Dr. B. Lewis and Mr. B. Karlovitz and on careful evaluations of plant geometries, current literature and accident parameters, that the existing experimental data on transition from de-flagration to detonation do not support as credible the possi-bility of its occurrence in an ice-condenser plant. It is within this reference frame that the Whiteshell pipe-sphere test configuration was conceived, that is, with the intention of investigating flame acceleration and propagation in general.
The modeling of an ice-condenser geometry for transition from deflagration to detonation phenomenon was found to be complex and impractical. There is a general consensus in the technical community that data collected from reduced scale models or models which do not reproduce the exact configuration, including material surface roughness, initial temperature, etc., of an ice condenser would not be conclusive. In view of this situation, a commitment was made to utilize the pipe-sphere configuration at Whiteshell as a best effort to acquire more fundamental knowledge on the phenomenon of transition to detonation and not to model the ice-condenser geometry specifically.
(d) As part of the experimental effort at Whiteshell, described in the response to Question Nll(c), AEP, TVA and Duke have undertaken to study 'the effects of physical phenomena which
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result from the presence of obstacles in the event of hydrogen combustion. A series of experiments are being concluded in which effects of turbulence generated by, either a fan or the presence of gratings are being investigated.
It is our belief that under conditions which are prototypic of an S2D type event, the Whiteshell turbulent experiments wa.ll adequately address conditions similar to those in an ice-condenser plant. Moreover, in calculations pertinent to the hydrogen mitigation and control system, conservative assumptj:ons'ave consistently been applied which would cor-respond to the existence of turbulent effects upon the overall
,safety evaluation of the plant. The forthcoming results from Whiteshell should confirm the validity of the assumptions used.
(e) The acceleration of a flame is generally recognized as a complex phenomenon. It is known that rapid flame acceleration requires strong turbulence. which may be generated by interaction of flow ahead of the flame with interior surfaces and obstacles.
The turbulence generated in a given situati;on and its effect on the flame acceleration process is strongly dependent on the geometry and othe factors such as heat transfer in the regi'on 'of i'nterest, There are no theories at present for quantitatzye prediction of flame acceleration, Experimental "
results obtained for a certain geometry are applicable to that geometry alone and cannot be extrapolated to other geometries.
In particular, Lewis and Karlovitz have stated that "the L/D obtained from experiments in pipes is applicable only to pipes and xs not applicable to the ice condenser".
Moreover, the geomet'ric ratio, L/D, is not necessarily a representative constant to be used to characterize the ice bed flow channels since they differ from a circular pipe for the purpose of flame propagation. The geometry of the ice bed is far less constrictive than that of a closed pipe.
The flow channel surrounded by each set of four ice baskets is actually open on all four sides between the baskets over their entire height. The melting of ice as the accident progresses would only open up the ice basket flow paths even more. Therefore, the sideways confinement in the ice bed is limited and is unlike experiments in pipes with complete radial confinement and one-dimensional flame propagation.
Furthermore, operation of the distributed ignition system will maintain the hydrogen concentration in the lower compartment at, or below 8 v/o. Burning at such low concentrations would produce weak flames with very little potential for flame acceleration. Hence, based on the lack of one-dimensional confinement in the ice bed flow channels and the maintenance of low concentrations of hydrogen, we believe that combustion initiated in the lower compartment and propagating through the ice condenser is not expected to be substantially different from that occurring in other parts of the containment.
Item 12 Information relevant to the consequences of prolonged combustion in the ice bed or upper plenum region is contained in Attachment No.
1 of Reference 5. The Duke Power Company (Duke) has performed a detailed three dimensional analysis of the heat-up of the ice con-denser wall panels when subjected to continuous combustion (in the ice bed) for forty-five minutes. The results of the Duke analysis, presented in Section 5.5 of Volume .3 of Reference 6, clearly indicate that the maximum insulation temperature (253 F) remained well below that at which pyrolysis of the foam would begin to occur. The maximum temperature of the sheet metal in contact with the foam insulation reported by Duke was 350 F.
As noted in Attachment No.' of Reference 5, fiber'glass encapsulated in. polyethylene baga is employed in the Cook Units for insulation purposes, whereas the 'Sequoyah and McGuire plants utilize polyeurethane foam. This design difference not, withstanding, we believe the aforementioned Duke analysis to be conservatively applicable to Cook Plant.
Item 13 A discussion of the effects of continuous combustion on the upper plenum~i nitionassemblies is contained in Attachment No. 2 of Reference 7 (See response to DCC Items A.l.a and A.l.b). Additional information on this topic is contained in Section 5.4.1 of Reference 6.
Item '14 The requested information will be transmitted at a later date along with the responses to items 7 and 8.
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Re'fe'r'ence'sfo'r Attachm'en't No'.1O'EP:NRC:0'0'5'OOG
- 1) Letter No. AEP:NRC:00500C dated 29 May 1981.
- 2) Letter dated Ol December 1981, L. M. Mills (TVA) to E. Adensam (NRC).
- 3) TVA Report, "Seguoyah Nuclear Plant-Containment Response to Degraded Core Events", 16 November 1981.
- 4) OPS Documents No. OPS-07A35, "The CLASIX Computer Program for the Analysis of Reactor Plant Containment Response to Hydrogen Release and Deflagration," G. M. Fuls.
- 5) Letter No. AEP:NRC:00500E dated 02 July 1981.
- 6) Duke Power Company Report, "An Analysis of Hydrogen Control Measures at. McGuire Nuclear Station", three vcr.urges,October 1981.
- 7) Letter No. AEP-NRC:00500F dated 28 October 1981
Attachment No. 2 to AEP:NRC:00500G Donald C. Cook Nuclear Plant Unit Nos. 1 and 2 Additional information on Hydrogen Mitigation and Control Responses to Mr. S. A. Varga's letter of 04 September 1981
The requested information is contained in Section 2.2 of the TVA Equipment Survivability Report (ESR) attached to their letter of Ol December 1981 to the NRC (Reference 1). We have reviewed the ESR and concluded that the analyses presented in Section 2.2 are conservative for application to the Cook Plant. It should be noted that the temperature profiles used for the lower compartment analysis do not account for the heat removal capacity of lower containment sprays and hence are conservative for application to the Cook Plant.
SNP Items A.3 throu h A.S The requested information is contained in Sections 2.2 and 2.3 of the ESR attached to Reference 1.
DCC Items I
Responses to questions addressed to the Cook Plant dockets are contained in Attachment No. 2 to AEP:NRC:00500P (Reference 2).
References for Attachment No. 2 to AEP:NRC:00500G
- 2) Letter No. AEP:NRC:00500F dated 28 October 1981.
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Attachment No. 3 to AEP:NRC:00500G Donald C. Cook, Nuclear Plant Unit, Nos. l and 2 Additional Information on Hydrogen 1Iitigation and Control Report on Hydrogen Combustion Control Studies
I EFFECT OF IGNITOR LOCATION AND WATER FOGS ON HYDROGEN COMBUSTION WITHIN AN ENCLOSED COMPARTMENT PROJECT REPORT December, 1981 Prepared by:
F. G. Hudson, Duke Power Company K. K. Shi'u, American Electric Power R. C. Torok; Acurex Corporation .
J. J. Wilder; Tennessee Valley Authority Project Conducted by:
Acurex Corporation 485 Clyde Avenue Mountain View, California Project Sponsors:
American Electric Power Service Corp.
Duke Power Company Electric Power Research Institute Tennessee Valley Authority
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TABLE OF CONTENTS 1.0 Introduction 2.0 Test Facility 3.0 Test Matrix and Procedures 4.0 Test Results 5.0 Conclusions Appendix Gas Chromatography Analysis
Section 1 Introduction Approximately ten hours into the accident at Three Mile Island, a hydrogen burn occurred inside the containment. Although this burn posed no real threat to the TMI containment, it'id create interest in hydrogen combustion and its effects on containment structures. The operating license applications for McGuire and Sequoyah Nuclear Stations contributed to the growth of this interest into a major safety concern, especially for ice condenser containments. The individual and joint activities of the three utilities owning ice condenser stations (American Electric Power, Ouke Power Company, and the Tennessee Valley Authority) are well documented in various licensing submittals,.or licensing proceedings and will not be repeated here. However, when the three utilities decided to install a distributed ignition system as a hydrogen mitigation system, the question of ignitor location within a compartment arose. Additionally, concurrent with the design of a distributed ignition system, several independent organizations suggested coupling a water fog system with the distributed ignition system to act as a pressure suppressant during combustion. To investigate the effect of ignitor locati'on on hydrogen combustion within a compartment, the three utilities', in conjunction with the Electric Power Research Institute, con-tracted with Acurex Corporation to conduct a series of tests. These tests were conducted at the SRI International Explosives Test Site near Livermore, California.
Although analyses had shown a pressure suppressant was not necessary for ice condenser containments, the utilities believed that investigating the effect of water fogs on hydrogen combustion could be of some potential interest to the
industry. Therefore, an investigation of water fog effects on hydrogen combustion was added to the Acurex project. This report presents the results of that project.
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Section 2 Test Facility 2 1
~ TEST VESSEL ANO MECHANICAL SYSTEMS The test vessel selected for this project has a volume t
of approximately 630 fthm. Vessel dimensions are: internal diameter - 7 ft., overall height - 21 ft., and "barrel" height - 17 ft. Auxiliary mechanical systems provide the ability to: 1) inject hydrogen or a hydrogen/steam mixture into the lower portion of the test vessel, 2) supply a water spray or microfog from the upper portion of the test vessel, 3) obtain pre-test and post-test vessel atmosphere samples, and 4) provide a means of premixing the vessel atmosphere for quiescent tests. Additionally, the capability to ignite the vessel atmosphere from the top, middle or bottom of the vessel was provided. A schematic of the test vessel and its auxiliary mechanical systems is presented in Figure 2-1.
A propane-fueled boiler supplied steam for the facility. This steam served as a parameter for several tests, as well as to preheat the test vessel to the desired temperature. The steam flowrate was monitored with an annular flow sensor and a differential pressure gauge. When steam was not required as a test parameter,'he boiler was isolated from the test vessel after preheating was completed. Bottled hydrogen served as the hydrogen source for the test vessel. The hydrogen flowrate was monitored with a rotameter and controlled with a control valve.
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A three horsepower electric motor and gear pump supplied water to the test vessel spray nozzles. A bypass loop was included to control the flowrate.
Utilizing a closed loop spray system, i.e., recirculating the spray water, avoided the potential problems associated with accumulating large volumes of water within the vessel. For the Phase 1 tests, a single Sprayco 1713 nozzle, 15 gpm flowrate, was mounted at the top of the test vessel. A manifold containing nine Sprayco 2163-7604 pinjet nozzles was mounted at the top of the test vessel for the Phase 2 tests. Oepending on the pressure drop across the nozzles, the total spray flowrate for the Phase 2 tests varied from l. 1 to 1.4 gpm. The spray manifold was constructed so 'as to provide an even spray distribution throughout the test vessel.
An air-operated fan was mounted inside the test vessel to assure a well mixed vessel atmosphere prior to the quiescent tests. Use of an air-operated fan eliminated the potential of an electrical malfunction resulting in a spurious ignition. The fan's air exhaust was vented outside the test vessel to avoid diluting the vessel atmosphere.
Two 4 inch butterfly valves located at the top and near the bottom of the test vessel allowed the vessel to be purged following the completion of each test.
A squirrel-cage" blower attached to the lower butterfly valve provided the motive force for purging the vessel. The vessel contents were vented to the atmosphere through the upper butterfly valve. The vessel was not vented until post-test samples were obtained.
7' Vessel atmosphere sample taps were located near the top and near the bottom of the test vessel. A remotely operated solenoid valve isolated each of the two sample lines from the test vessel. Mhen a solenoid valve was open, the vessel atmosphere sample was pumped through a cold trap to remove water. The sample then passed through a silica gel trap to remove any remaining moisture. The sample then flowed,through a gas meter into a glass sample bottle. A sample was extracted from the sample bottle by a syringe and injected into a gas chromatograph.
(A detailed discussion of the gas analysis methodology is presented in Appendix A).
Two ignitor assemblies supplied by Duke Power Company were mounted inside the test vessel. The ignitors were located on the vessel centerline at either the top, middle or near the bottom of the test vessel. Only two ignitor locations were occupied at one time. The top ignitor location was not used during the Phase 1 tests requiring sprays or during any of the Phase 2 tests since an ignitor assembly located at the top effectively created a significant spray/fog maldistribution within the test vessel.
2.2 INSTRUMENTATION The test vessel was instrumented to provide the following information: vessel atmosphere temperature, vessel wall temperature, flame front propagation, and vessel pressure. Three mil Type K thermocouples were used to measure temperatures and detect flame front propagation. Strain gauge pressure transducers and piezoelectric pressure transducers were used to measure vessel pressures. A schematic of the
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test vessel instrumentation is presented in Figure 2-2.
t As just mentioned, 3 mil thermocouples were used to measure temperatures and detect flame front propagation. Hale Type K thermocouple jacks were used as attachment points for the 3 mi 1 thermocouple junctions to increase the robust-ness of the thermocouple. The junctions were located between .the jacks with the leads attached directly to the jacks. Vessel wall temperature thermo-couples were welded directly to the vessel wall.
To determine the flame front propagation pattern within the vessel, a special electronic circuit was developed. This circuit used a high input impedance operational amplifier comparator to detect the temperature rise in the 3 mil Type K thermocouples located within the vessel. A schematic diagram of the flame front detector circuit is shown in Figure 2-3. Five circuits were used, each circuit monitoring seven thermocouples located in a vertical array within the vessel. The 5 x 7 grid was located vertically on a plane formed by the diameter and centerline of the test vessel as shown in Figure 2-2.
Mhen one of the thermocouples in the grid was exposed to the flame front, the thermocouple output voltage would rise and trigger the comparator. This, in turn, placed a signal at one input point of the digital-to-analog converter.
The output from the OAC was then recorded. Since the input signals from the thermocouples-could be considered as binary bits, every voltage that was generated by the OAC corresponded to a discrete combination of hot thermocouples.
Thus, by comparing the output voltage against time, the .instant that the flame front arrived at each location could be determined, and a map derived showing flame front propagation.
Two types of pressure transducers were used. Bell and Howell CEC Model 1000 strain gauge pressure transducers were used for static and slow response conditions. These were powered by CEC 1-183 strain gauge signal conditioners located within a CEC 1-080 power supply chassis. To record high frequency transient pressure pulses, PCB Piezoelectronics Model lllA24 piezoelectric pressure transducers were used. These transducers were powered by a Model 484B10 power supply.
Two recording systems were used for data acquisition. A twenty eight channel FM tape recorder, EMI Model 7000C, was used to record all potentially fast response data. Frequency response of the unit was 10kHz or greater. An Autodata 9 datalogger. recorded relatively slow response signals and served as a backup to the FM tape recorder.
l f1WRE 2-1 TEST FACILITY f I ECHAN I CAL SCHEf 1ATI C Blast all -
Propane Vent Control room Boiler Gas Spray or fog sampling Ignitor n 0
N Blonde I Rotameter C Pump 14 ter supply r igni tor Gas fan sampling Cxhdust Air Drain
FIGURE 2-2 TEST FACILITY INSTRUMENTATION SCHEMATIC F Flame front gage T Thrrmocouple wT Mall thr~couple SXO SXO Strain gauge pressure transducer Pxo Pirxoelectrlc pressure transducer rxo STG Strain gauge SG Silica grl ~ate( trap STG 0(W Ory gas meter GO Glass bottle Ignltor GC Gas chromatograph F F F F f T
OGH SG Pump F F F F F GG F F F F F T
F F F F lgnltor OGli, SG S lc Pur ip F F F F' T
lgnl tor F F F F Ml'. Sxo F F F F rxo T
Sxo Injrc t ion nol l l e
~ 8
I
~ ~ 1
~g FIGURE 2-3 FLAME FRONT OETECTOR CIRCUIT .
415Y 4
301 THIRtSOCOVPlE I.ID 3290 )7 45 45 44 43 I2 II out 4I 20 0 0 0 0 0 0 I 0.0)SV 0 0 0 0 0 I 0 0. I 5 2 V 0 I 0. 314V 0'00 0 0
0 I 0 0
0 0
0 0. 621V 3290 0 0 I 0 0 0 0 1.255v 41 0 I 0 0 0 0 0 2.510v I 0 0 0 0 0 0 5.02ov 0 0 0 0 0 I I 5.099V IIC I I I I I I I IOY 3290 42 I4pe Drct 0-lov Output 3290 I~
42 v 15V 3290 e3 Ana ln9 Ppv>CPS AI)558 OAQ 3290 43 3290 i4
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Section 3 Test Matrix and Procedures 3.1 TEST MATRIX The test program was designed to investigate the effects of ignitor location and water fogs/sprays on hydrogen combustion. Test parameters were selected based on degraded core analyses conducted by American Electric Power, Duke Power Company, and TYA. Flowrates for steam and hydrogen were scaled by the ratio of test vessel volume to the combined lower compartment and deadended compartment volumes of an ice condenser containment. These scaled flowrates were derived from the average steam and maximum hydrogen rel'ease rates dur-ing the hydrogen generation portion of an S~ 0 accident sequence. Containment response analyses conducted by the aforementioned utilities indicated that 160'F was a maximum atmospheric temperature that would be encountered in the lower and deadended compartments at the onset of hydrogen generation. Hy'drogen or hydrogen/steam mixtures were injected into the lower portion of the test vessel since many subcompartments within an .ice condenser containment are accessed from the bottom.
Spray nozzle fl'owrates were determined by the desired mean droplet diameter.
I The Sprayco 1713 spray nozzle, a model used in containment spray systems, was operated at a GP of 40 psi. This provided a flowrate of 15 gpm and is the designed operating condition in a containment spray application. Vendor supplied information indicated that the number mean droplet diameter at this flowrate is 200'. Based on studies conducted by Factory Mutual Research
Corporation, the Sprayco 2163-7604 nozzle was selected for use in the water fog tests. Heasurements taken at Factory Hutual Research Corporation'ndicated that when operated at hP's of 20 psi and 30 psi the resulting number mean droplet diameters were lip and Sp, respectively. The total fog flowrates from the test vessel's nine nozzle manifold were 1. 1 gpm and 1.4 gpm, respectively.
Tests outlined in Table 3-1 were intended to investigate the effect of ignitor location on combustion. Both hydrogen and hydrogen/steam mixtures were injected to account for the potential variation in the transient conditions within an
~ ice conde'nser containment. Tests with sprays were included to account for the presence of subcompartment sprays in some containment designs. As mentioned earlier, the steam and hydrogen flowrates of 2. 1 and 0.035 lbmlmin were based on transient analyses. Tests 1.6 and 1.7 were conducted with the hydrogen flowrate arbitrarily increased by a factor of three. Test 1. 11 was conducted to observe the effect of a lower preheat on combustion characteristics. The duration of tests was determined by the time required to obtain the same relative hydrogen mass injected in the test vessel as is calculated in the previously mentioned degraded core analyses. All tests were conducted with the test vessel atmosphere initially saturated.
Zalosh, Robert G., "Mater Fog Inerting of Hydrogen Air Hixtures," Factory Hutual Research Corporation, September, 1981.
\
1 I
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~
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The water fog test matrix is presented in Table 3-2. quiescent tests were conducted as a basis of comparison to observed results of the dynamic tests.
The basis for all remaining test parameters was discussed above.
3.2 TEST PROCEDURES Two types of tests were conducted: quiescent and dynamic. A known amount of hydrogen was injected into the test vessel for the quiescent test prior to energizing the ignitor assembly. The dynamic tests consisted of injecting hydrogen or hydrogen/steam -into the vessel with the ignitor assembly pre-energized.
For all tests, the test vessel was pre-heated to the desired temperature and the instrumentation and data acquisition system was checked and ca1ibrated.
After the completion of each test, the test vessel-fan was turned on and a post-test sample obtained. Subsequently, the test vessel was purged.
Test procedures varied slightly for the quiescent and dynamic tests. For the quiescent tests, the vessel fan was turned on after completing the pre-test activities mentioned above. A known amount of hydrogen was injected into the vessel and a pre-test sample. was obtained. The vessel fan was then turned off and, if required, vessel sprays actuated. At this point, the data acquisition system and the'ignitor assembly were energized. For the dynamic tests, the ignitor was energized after completion of pre-test activities. If necessary, vessel sprays were then actuated. The data acquisition system was energized prior to the initiation of hydrogen injection.
TABLE 3-1 Ignitor Location Test Natrix" Ignitor Hydrogen F1ow (ibm/min) Steam Flow (ibm/min) Spray Flow Test Location 0.035 ~
0.105 2.1 ~15 m Top ' X 1.2 Top 1.3 Bottom 1.4 Bottom 1~ 5 Bottom
.1. 6 Bottom 1.7 Bottom 1.8 Center 1.9 Center X
- 1. 10 Center
- l. 11 Bottom "Test vessel was preheated to 160'F for all tests except 1. 11. Test l. 11 was performed with 120'F preheat.
TABlE 3-2 Water Fog Test Matrix*
Hydrogen Hydrogen Flow (ibm/min) Steam Flow (ibm/min) ,
Fog Nozzle Pressure (psi)
Test ¹ v/o 0. 035 0. 105 2.1 20 30 2.1 5.0 2.2 7 5
~
2.3 10. 7 2.4 10. 7 2.5 10. 7 2.6 7.5 2.7 7.5 2.8 X 2.9
- 2. 10""
- 2. 11
- 2. 12
- 2. 13 "Test vessel was preheated to 160'F for all tests. Ignitor located near the bottom.
"*Vessel mixing fan was operating.
Section 4 Test Results OEFLAGRATION CHARACTERISTICS Based primarily on pressure and flame front detector data, two distinct types of deflagrations occurred. These deflagrations were termed "discrete" and "intermittent". A "discrete" deflagration was characterized by a rapid pressure and temperature rise. The duration of the burn appeared to be dependent on the fraction of the vessel volume that could support a propagating hydrogen flame.
By comparison, an "intermittent" deflagration appeared as repeated burns accompanied by much slower and lower pressure/temperature rises. Observed deflagrations were further categorized as "major" or "minor". A "major" deflagration, whether discrete or intermittent, occurred throughout the test vessel. A "minor" deflagration, on the other hand, was localized in nature.
Pressure and temperature histories typical of a major-discrete deflagration are presented in Figure 4-14. "CENTERLINE Tl" and "CENTERLINE T4" are centerline thermocouples located near the top and bottom of the test vessel. "VESSEL PRESSURE 2" is a strain gauge pressure transducer located in the lower portion of the test ves'sel. As seen in the figure, the test mixture ignited approxi-mately twenty seconds into the test. The periodic disturbances in the two temperature traces were caused by the datalogger scanning these channels. The flame front detector output for the same test is presented in Figure 4-15. The five data channels correspond to the five vertical columns of transducers shown
'n the inset figure; each channel receiving the output of seven flame detectors.
The size of each step is determined by the combination of detectors that triggered. A fullscale reading indicates that all seven detectors within a channel have triggered. Of the channel outputs presented in Figure 4-15, only channel 5 was not reading fullscale. The flame front detector data indicated that approximately 95K of the flame front detectors triggered. This, therefore, was classified a "major" and "discrete" deflagration.
Figures 4-5 and 4-6 present temperature, pressure, and flame front detector data y
typical of major intermittent deflagrations. The flame front detector data in Figure 4-6 indicate that all channels were triggered, signifying that a major deflagration occurred. Temperature data in Figure 4-5 shows a sharp, but not extremely large, temperature increase that remained relatively constant for several minutes before starting to slowly decay. This indicates that intermit-tent deflagrations were occurring. Note that near the end of this, particular test, a major discrete burn occurred.
Minor intermittent burning is demonstrated by the temperature, pressure and flame front detector data presented in Figures 4-21 and 4-22. Flame front detector data, Figure 4-22, 'indicate that few flame detectors were triggered during this test. This characterizes a "minor" deflagration. The temperature and pressure data in Figure 4-21 show the small and gradual increases characteristic of intermittent burning. Figure 4-3 also provides. temperature and pressure histories indicating minor intermittent deflagrations. Note that near the end of this test, several minor discrete deflagrations occurred.
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4.2 INSTRUMENTATION UNCERTAINTY Test data was primarily test vessel pressure/temperature histories and flame front detector output. Additionally, vessel atmosphere constituents were deter-mined via a gas chromatograph. To properly evaluate the test data, it was necessary to know the errors associated with the instrumentati'on used.
Sources of error in the thermocouple data were: thermocouple material and 3unction uncertainty, thermocouple amplifier error, test facility offset errors due to electrical ground loops, tape recorder input and playback error, analog to digital conversion errors and plotter inaccuracy. Standard Type K thermo-couple error estimates were +2.2 C from 0 to 278'C and +3/4X above 278'C, ANSI Standard C96. 1. The low temperature range error corresponded to approximately
+3K when peak temperatures were around 100'C. The estimated gain error for each of the three signal amplifiers was +EX. Errors associated with the analog to digital conversion and plotter inaccuracies were considered negligible.
Therefore, using the root-mean-square method, the total random, error for low temperature cases was approximately, 1
[3> + 12 +1> +1>]~ = 3. 5$
and for high temperature cases,
[(3/4) +1 + 1 + 1 ] = 1.9X.
I C
Pressure data were obtained from two strain gauge pressure transducers mounted at the top and near the bottom of the vessel. Although data from the piezoelectric pressure transducers were recorded, the observed pressure transients did not warrant use of the piezoelectric instead of the strain gauge pressure transducers. Sources of error in the pressure data were: transducer error, the transducer amplifier error, and the same tape recorder amplifier errors encountered in the thermocouple data. The manufacturer's estimated error for the transducer was +.25%%d of full range. This corresponded to +0.5/ of the signal resulting from a large hP, i.e., a major deflagration, and +2.5%%u'or a small hP. Using the root-mean-square method, the total random error for high hP cases was approximately, 1
I'(q)> + 1~ + 1~ + 1~]~ =,]..g) and for low AP cases, 1
[2.5 + 1 + 1 + 1 ] = 3.0X.
Calibration and test gas chromatograms were analyzed using the peak height method. Water vapor corrections were made to convert dry gas sample analyses to actual test'vessel conditions.
Uncertainty estimates for the resulting gas constituent volume percentages were based on the repeatability of sample analyses. Normalized mean peak heights were calculated from the analyses of each test run. The ratio of the largest for each test run estimate
~
deviation from the mean to the mean was used as an
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e of the gas analysis uncertainty. The mean estimated uncertainties were approximately +lOX for the hydrogen analysis and +20K for the oxygen analysis.
Problems encountered with the gas chromatograph's thermal conductivity detector were suspected as being a primary cause of scatter in the data. Oiscussions with the manufacturer indicated that filament oxidation was hampering performance.
4.3 IGNITOR LOCATION TEST SERIES Tests were conducted by varying the ignitor location in three test environments:
hydrogen injection, hydrogen/steam injection, and hydrogen/steam injection with sprays. Two additional tests were conducted with the hydrogen flowrate arbitrarily increased by a factor'of three. The final test of this series was conducted with a reduced vessel pre-heat. A summary of the results obtained from this test series is presented in Tables 4-1 and 4-2.
Figures 4-2, 4-4, and 4-10 provide the pressure histories 'from ignitor location tests without steam or spray (tests 1.2, 1.5, and 1. 10). At lean hydrogen con-centrations, flame propagation is only upwards. With this in mind, it was anti-cipated that while localized burning was occurring in the vicinity of the top ignitor, the hydrogen concentration would increase throughout the remainder of the vessel. When the flammability limit for downward propagation was reached, a major discrete deflagration would occur. This appeared to be the sequence of events in test 1.2 with minor intermittent deflagrations beginning a't 300 seconds followed by a major discrete deflagration at 580 seconds. The maximum pressure
. rise was expected to be smaller for the center ignitor location (test 1. 10) than the top location because of the increased vessel volume that would be
(W exposed to upward propagating flames at lean concentrations. Table 4-1 shows that the pressure rise was lower by approximately a factor of three. Hinor intermittent deflagrations began around 220 seconds and continued throughout the test. The lowest ignitor location was expected to produce an even milder pressure rise since a substantial portion of the vessel would be exposed to upward propagating flames. However, as Figure 4-4 shows, that was not the case. Apparently, the relative locations of the injection port and the lowest ignitor location precluded the ignitor from igniting hydrogen early in the test.
The major discrete deflagration that occurred at 4SO seconds indicated that the injection flow apparently bypassed the ignitor until most of the vessel contained a flammable mixture. The resulting deflagration produced a higher pressure rise than that attained by the top ignitor for two apparent reasons: 1) The top ignition was preceeded by localized deflagrations; thus reducing the mass of hydrogen within the vessel; 2) Flames propagate slower downward than upward; thus allowing more time for heat transfer.
Figures 4-1, 4-3, and 4-9 show that results from the tests with hydrogen/steam injection (tests l. 1, 1.4, 1.9) were similar to those obtained from the hydrogen injection tests. It was anticipated that the hydrogen/steam injection tests woul,d yield milder pressure increases. This was due to steam impeding, the combustion process as well as acting as a diluent; thus reducing the flame propagation velocity. This would result in increased heat transfer and decreased temperatures/pressures. Additionally, adding steam increased the injection velocity from approximately 0.7 ft./sec. to 5.7 ft./sec. This was believed to increase mixing within the vessel and thus allow deflagrations to occur at leaner
hydrogen concentrations. Table 4-1 indicates that the peak pressures at the three ignitor locations were reduced with steam added to the injection flow.
The most dramatic change was with the bottom ignitor"location. Apparently, the increased mixing provided by the steam flow allowed the lowest ignitor to '
function as discussed previously. The top ignitor provided the largest pressure rise, with the center and bottom ignitors being approximately a factor of three less.
The addition of a water spray was expected to create some amount of turbulence within the vessel that would enhance mixing and allow combustion to occur at leaner hydrogen concentrations. It was also anticipated that a water spray would act as a dispersed heat sink; thus further reducing temperatures and pressures, Table 4-1 shows that the addition of water spray, tests 1.3 and 1.8, c did reduce the maximum pressure. The bottom ignitor yielded only a very slight pressure rise with no corresponding flame front detector activity. However, post-test atmosphere analysis indicated that combustion had occur red. These deflagrations must have been very localized near, the ignitor and apparently relied upon spray induced turbulence for a continual supply of lean hydrogen mixtures. Figure 4-8 shows that the center ignitor provided a series of minor discrete burns. The pressure rise was slightly higher than obtained from the bottom ignitor." Test vessel design did not allow a spray test to be conducted with the upper ignitor location.
Two tests were conducted with the hydrogen flowrate arbitrarily increased by a factor of three. ~ One test was conducted with hydrogen injection, test l. 7, and the second with hydrogen/steam injection, test 1.6.. The bottom ignitor was used
for both tests. Comparing tests 1.7 with 1.5 and 1.6 with 1.4 shows that the transients were similar to their low flow counterparts, with the exception of ignition occurring ear lier in the tr ansient. In test 1. 7, ignition occurred slightly earlier than the 150 second ignition expected from a higher flow rate (see Figure 4.7). It is possible that the increased injection velocity had a slight effect on mixing within the vessel. This would allow an ignitable mixture to reach the ignitor earlier. This would also explain the slightly lower pressure ri-se from test 1.7 since less hydrogen would be present within the vessel at ignition. Note that the high flowrate pressure rise was 85K of the low flowrate pressure rise and that the high flowrate ignition time was 85K of the anticipated 150 second ignition time. Adding steam to the high hydrogen flowrate, test 1.6, yielded deflagrations similar to the low flowrate counterpart, test 1.4, but a pressure rise essentially identical to that obtained from test 1.7. Figure 4-5 shows that ignition occurred at approximately 100 seconds, one third of the 300 second ignition time for test 1.4. The ensuing intermittent deflagrations were more severe in test 1.6 because the higher hydrogen flowrate apparently resulted in a higher energy release rate. Why these intermittent deflagrations were not followed by repeated discrete deflagrations as seen in test 1.4 is uncertain. One possible explanation is that with vessel atmosphere temperatures in excess of 400'F for over one third the duration of test 1.6, a fraction of the water collected at the tank bottom from vessel pre-heating was vaporized during the intermittent burning. This could have caused the defla-grations to be very localized, similar to those obtained in test 1.3. Thus, hydrogen could have built up in the vessel while the steam was slowly condensing
. until an ignitable mixture was once again obtained. The result would be a lull in flame front detector activity followed by a major discrete deflagration.
V Figures 4-5 and 4-6 show such characteristics. Some credence is lent to this possibility by noting that the post-test water concentration from test 1.6 was 50K larger than that obtained from test 1.4.
One test was conducted, l. 11, with the vessel pre-heat reduced to 120'F from 160'F. The results obtained were very similar to those obtained from 1.3, an identical test with a 160'F vessel pre-heat. A very slight pressure rise occurred with no corresponding flame front detector activity observed. This indicated that the burn, as in test 1. 3, was very localized.
4.4 WATER FOG TEST SERIES Tests were conducted to investigate the effects of a water fog on hydrogen combustion. The fog nozzle, Sprayco Model 2163-7604, created different fog characteristics depending on the pressure drop across the nozzle. Tests were conducted with two different fogs. Based on data obtained from Factory Mutual Research Corporation, a 20 psi bP yielded a fog with a number mean droplet diameter of lip, while a 30 psi bP yielded a number mean droplet diameter of 8p. Two types of tests were conducted: quiescent and dynamic. The dynamic tests were conducted with arid without steam. All tests utilized the bottom ignitor. A sum'mary of the results obtained from this series of test is presented in Tables 4-3 and 4-4.
I 4.4.1 QUIESCENT TEST SERIES To provide a baseline of information for evaluating the dynamic fog tests, a series of quiescent tests were conducted. Nominal hydrogen concentrations of 5, 7.5, and 10K were selected. Table 4-4 indicates that the completeness of combustion for-those tests without fogs (tests 2. 1, 2.2, and 2.3) was approximately 30K, 90K, and 99K, respectively. This data agrees reasonably well with published data. The temperature and pressure histories of these three tests are presented in Figures 4-11, 4-12, and 4-13. Further tests with 5X hydrogen-were not con-ducted.
Repeating these tests with fogs present, a significant decrease in pressure rise was anticipated. Oue to .the large surface area present within a fog, .the fog was expected to act as a dispersed heat sink; resulting in reduced temperatures and pressures. However, as Table 4-3 indicates for the 7.5X hydrogen tests, 2.6 and 2.7, the observed pressure rises were slightly higher. Table 4-4 shows that the completeness of combustion increased from approximately 90 to greater than 99K. This indicated that these particular fogs acted Very much like sprays for lean hydrogen concentrations. The turbulence created by the fog flow apparently enhanced the completeness of combustion; thus increasing the pressure rise. The heat sink effect of the fogs was evident from the slightly lower temperatures (compare Figure 4-12 with Figures 4-17 and 4-18). The fogs had no apparent effect on the peak pressure rise in the 10K tests, 2.4 and 2.5.
4.4. 2 DYNAMIC TEST SERIES Hydrogen injection tests 2.8 and 2. 12 were identical to test 1.5 except that fogs were included. Results from 2.8 and 2. 12 indicated only minor intermittent deflagrations (see Figure 4-23). The observed pressure rise was an order of magnitude lower than that observed in test 1.5. It would be reasonable to assume that a great deal of the pressure reduction was due to fog flow induced mixing, which allowed a flammable mixture to reach the ignitor earlier. Note that the effect of adding steam to test 1.5 (test 1.4) was minor intermittent burning early in the transient with a pressure rise of 2.3 psi; the same effect observed in tests 2.8 and 2. 12 (see Figure 4-3).
Figure 4-19 shows a pressure history typical of that obtained from hydrogen/steam injection tests with fog present, tests 2.9 and 2. 13. The observed pressure rises were similar to that obtained in test 1.4, hydrogen/steam injection with no spray. In both cases, with and without fog present, ignition occurred at approximately 300 seconds. However, in test 1.4 the result was minor intermit-tent deflagrations eventually becoming minor discrete deflagrations. Tests 2.9 and 2. 13 provided minor discrete burns immediately upon ignition. As a res'ult, it appears that more hydrogen was consumed in the fog tests than in the non-fog tests. This would also appear to,indicate that a major contribution from the generation of fog in this test series was to provide uniform mixing within the vessel.
One test was conducted with the hydrogen flowrate arbitrarily increased by a factor of three in the presence of a fog, test 2. 11. Figures 4-21 and 4-22 show
the pressure history and flame front detector activity obtained from this test.
This test provided an indication of the heat sink effect of a fog as well as its effect as a source of turbulence. Note on Figure 4-21 that ignition occurred approximately 20 seconds into the test, while in test 1.7, an identical test without fog, Figure 4-7 shows that ignition did not occur until 130 seconds into the test. This difference in ignition time could be accounted for by the mixing created by the fog. The heat sink effect was apparently demonstrated since the peak temperature of test 2. 11 remained around 200'F, while the peak temperature in test 1.7 hovered around 400'F. In test 2. 11, fog produced turbulence apparently prevented a major discrete burn by inducing intermittent deflagrations early, thus precluding the relatively high hydrogen concentrations needed for a major discrete deflagration. Additionally, the fog acted as a heat sink during the resultant intermittent deflagrations; thus minimizing pressure and temperature increases.
F Comparing test 2. 11 with test 2. 12, an identical test with the lower hydrogen flowrate, the deflagration character-istics were similar with the high flowrate test having an earlier ignition.
Finally, data on the effect of fan induced turbulence was obtained when the mixing fan was accidently actuated prior to test 2. 10. The result was that ignition occurred earlier (see Figure 4-20) than in a similar test without the fan operating..('see Figure 4-23). Other than the ignition time, the results of both tests were relatively similar; minor intermittent deflagrations with ve'y slight pressure rises.
TABLE 4-1 Summary of Test Results: Ignitor Location Test Series Hax. hP Test ¹ Test Characteristics I nitor Location ~(si ) Oefla ration Characteristics 1.1 Low H2, steam Top 13 Minor, major intermittent 1.2 Low H2 Top 20 Minor intermittent, major discrete 1.3 Low H~, steam, spray Bot'tom 1 Minor intermittent 1.4 Low Hz, steam Bottom 4.5 Minor intermittent, minor discrete 1.5 Low H2 Bottom 28 Major discrete, minor intermittent 1.6 High Hz, steam Bottom 24 Major intermittent, major discrete 1.7 High Hz Bottom 23. 5 Major discrete, minor intermittent 1.8 Low H2, steam, spray Center 2.7 Minor discrete 1.9 Low H2, steam Center 4 Minor intermittent
- 1. 10 Low H2 Center 6 Minor intermittent
- 1. 11 Same as 1.3, lower preheat Bottom 1 Minor intermittent
TABLE 4-2 Test Vessel Atmosphere Constituents: Ignitor Location Test Series Test 0 Post-Test H2(v/o) HzO(v/o) 02(v/o) 1.1 11. 0 17. 6 8.9 1.2 2.6 23. 8 15. 1 1.3 6.5 19. 8 6.1 1.4 7.9 21. 0 5.7 1.5 2.1 38. 2 12. 2 1.6 10. 9 32. 1 6.5 1.7 12. 7 37. 6 1.7 1.8 7.9 30. 1 7.2 1.9 3.6 46. 1 5.2
- 1. 10 0.4 36. 3 5.7
- 1. 11 3.5 27. 8 2.6
I
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TABLE 4" 3 Summary of Test Results: Water Fog Test Series Max. DP Test I i 'Test Characteristics ~(sl) Defla ration Characteristics 2.1 quiescent, 5 v/o Hz 8 Minor discrete 2.2 quiescent, 7.5 v/o H2 36 Major discrete 2.3 quiescent, 10.7 v/o H2 48 Major discrete 2.4 quiescent, 10.7 v/o H2, fog 30 47 Major 'discrete 2.5 quiescent, 10.7 v/o Hz, fog 20 50 Major discrete 2.6 quiescent, 7.5 v/o H2, fog 20 40 Major discrete 2.7 quiescent, 7.5 v/o H2, fog 20 39 Major discrete 2.8 Dynamic, low H~, fog 20 2 Minor intermittent 2.9 Dynamic, low Hz, fog 20, steam 5 Minor discrete
- 2. 10 Dynamic, low H~, fog 30, fan 1 Minor intermittent
- 2. 11 Dynamic, high Hz, fog 30 2.9 Minor intermittent
- 2. 12 Dynamic, low H~, fog 30 1 Minor intermittent
- 2. 13 Dynamic, low Hz, steam, fog 30 1.6 Minor discrete
I 4
TABLE 4-4 Test Yessel Atmosphere Constituents: Mater Fog Test Series Test 8 Pre-Test Post Test H2(v(o) H20 (v/o) H2(v/o) H20(v/o) Oz(v/o) 2.1 4.7 25. 6 3.4 33. 2 ll. 5 2.2 7.8 26. 1 0.9 43. 2 10. 0 2.3 10. 2 24..9 <0.1 40. 4 10. 9 2.4 9.7 23 8
~ <0.1 44. 9 8.7 2.5 10. 2 25. 4 <0.1 40. 4 9.2 2.6 7.2 32. 9 <0.1 47. 6 8.4 2.7 7.6 29.4 <0.1 41. 6 11. 4 2.8 3.3 30. 1 6.8 2.9 6.2 45. 5 5.1
- 2. 10 2.8 49. 2 2.7
- 2. 11 13. 5 35. 1 <0 1
- 2. 12 <0.1 37. 0 9.2
- 2. 13 3.7 45. 4 3.7
(W FiGURE 4-1 TEST 1.1 VESSEL PRESSURE 1 60 400 50 300 i+ 4o C4 200 3O C4 V) D 40 20 l~
100 10 0 0 0 100 200 '300 400 500 600 700 000 900 1000 1100 1200 1300 TI.'.fE (Sl'.t:)
FIGURE 4-2 TEST 1.2 CENTERLINE T4:
1000 1600 800 1200 600 400 800 200 400 0
0 100 200 300 400 500 600 700 800 900 1000 1100 1200 1300 TIME (SEC)
VESSEL PRESSURE 1 60 400
-50 300 40 200 30 20
>00
-10 0- 0 0 100 200 300 400 500 600 700 800 900 l000 ll00 1200 1300 TT ~ i t. l'c-.I. f'3
F RE 4-3 TEST 1.4 CENTERLINE TP.
1000 1600 800 1200 600 400 800 200 400 t
0 100 200 300 400 500 600 700 800 900 1000 1100 1200 1300 TIME (SEC),
VESSEL PRESSURE 1 150 20 125 100 15 75 10 50 25 0 100 200 300 400 500 600 700 000 900 1000 1100 I'00 1300
.Vr ~~E (sr'.c)
FfGURE 4-4 TEST 1.5 CENTERLINE T4 100Q 1600 800 1200 600 400 800 20Q 400 0 100 200 300 400 500 600 700 800 900 1000 1100 1200 1300 TIME (SEC)
VESSEL PRESSURE 1 60 400 50 300 4Q 200 30 20 100 10 0 100 200 300 400 500 600 700 800 900 1000 1100 1200 1300 v r'~t'. (~t'r.'I
PHESSUI<E (KPA) TEMPERATURE (C)
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iURE 4-? TEST 1.7 CENTERLI NE .T2 000 1600 800 1200 600 400 800 200 400 0 0 100 200 300 400 500 600 700 800 900 TIME (SEC) VESSEL PRESSURE 1 60 400 50 300 40 200 30 20 100 10 0 0 100 200 ,300 400 500 600 700 800 900 T~,~r (~~>)
FIGURE 4-8 TEST 1.8 VESSEL PRESSUPE 2 150 20 125 100 15 C4 t~ 75 10 M cn 0 50 25
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FIGURE 4-9 TEST 1.9 i50 VESSEL PRESSURE I 20 125 100 i5 C4 A 75 D 10 D (/) (n 50 CO ~J 25 C4 0 -0 0 100 200 300 400 500 . 600 700 000 900 1000 1100 1200 1300 TI haft."(SL'C)
FIGURE 4- l 0 TEST 1.10 VESSEL PRESSURE 1 150 20 125 100 41 75 D 10 D 50 V) r~t 25 IX S4
,II 0 0 100 200 300 4 00 500 600 700 000 900 1000 1100 1200 1300 TitvfE (SEC)
GURE 4-11 TEST 2.1 CENTERLINE T4 1000 1600 800 1200 600 400 800 200 -~ ~
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10 20 30 40 50 60 70 80 90 100 TIME (SEC} VESSEL PRESSURE 1 150 20 125 100 r~ 75 10 if) 50 5 0 0 10 30 40 50 60 70 90 100 Tr WI'. (~rC}
F<6uRE O-12 TEST 2.2 CENTERLINE Tl 1000 1600 800 1200 600 400 000
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'ESSEL PRESSURE' 60 400 50 300 40 200 30 100 20 10 I
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FIGURE 4-13 TEST 2.3 CENTERLINE T2 1000 1600 800 1200 600
- 000 400 200 400 0
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I FIGURE 4-14 TEST 2.4 CENTERLINE Tl 1000 1600 800 1200 600
. 000 400 200 - 400 0-0 10 20 30 40 TIME (SEC)
. CENTERLINE T4 1000 1600 800 1200 600 400 000 200 - 400 0
10 20 30 40 TIME (SEC) VESSEL PRESSURE 2 60 400 50 300 40 200 30 U) 20 PJ 100 10 P 0 -0 0 10 20 30 0 TI~.(E (SEC)
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<0 200 30 20 CO 100 L 10 C4 0- -0 10 '20 30 0 TI! [E (SEC)
~ ~ FIGURE 4-32 TEST 2.6 CENTERLINE Tl 000 1600 000 1200 600
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CENTERLINE T4 1000 1600 800 1200 600 400 000 200 - <00 0, 10 20 30 40 TIME (SEC) VESSEL PRESSURE 2 60 400
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FIGURE 4-i8 TEST 2.7 CENTERLINE Tl 1600 800 C4'000 600 1200
- 000 S4 400 4l 200 - 400 10 20 30 40 TIME (SEC)
CENTERLINE T4 1000 160Q hl 800 C4 - 1200 D 600 9', - 000 400 C4 200 400 0-, 10 20 30 40 TIME (SEC) VESSEL PRESSURE 2 60 400 50 300 40 R 200 3Q D U) 20 CA 100 A' 10 0 10 20 <0 TIiifE (SEC)
FIGURE 4-19, TEST 2.9 VESSEL PRESSURE 2 20 125 ioo V) f/~ 50 haft'.
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FIGURE 4-20 FLAME FRONT ACTIVITY TEST 2.10 TL VAW'lW~ CHAt<HEI j. 12)il5 T~~WTM/ f f f f f CHAWi<EL 2 f f f f f f f f f f CHANNEL 5-f f f f CHAPliXEL 4 f f f f f f f f f f' f f / f CHANNEL 5
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Section 5 Conclusions
- 1. Location of an ignitor within the test vessel does affect the characteristics of hydrogen deflagrations. Lowering the ignitor location produces milder pressures during hydrogen combustion. This appears to be a result of increasing the fraction of the vessel volume exposed to upward propagating flames in lean hydrogen concentrations. The addition of steam and water sprays also reduced the pressure rise. However, the lower ignitor locations still produced milder pressures than the top location.
- 2. Fogs were thought to reduce the pressure rise resulting from hydrogen combustion. This was the case for dynamic tests, but not for quiescent tests. Mater fogs apparently enhance the rate of combustion. Thus, heat transfer is not as significant in quiescent tests causing the deflagration to be more like an adiabatic deflagration. For dynamic tests, water fogs promote mixing and allow ignition to occur earlier, resulting in lower energy release rates.
APPENDIX A Gas Chromatography Analysis Test vessel samples were obtained through nonheated '4 inch stainless steel probes located near. the top and bottom of the vessel (See Figure 2-1.). Each probe was connected to the inlet of a Thomas diaphragm pump. Vessel isolation was provided by solenoid valves, A ~4 inch stainless steel line connected the discharge of each pump to a stainless steel condenser coil submerged in an ice bath. Tarred silica gel columns were located at the outlet of the condenser coil to remove any remaining moisture. '4 inch stainless steel tubing carried the sample from the silica gel columns, through a dry gas meter, and into a 250
. mil glass bulb. Thermocouples monitored the inlet and outlet pressures of the , gas meter. Solenoid valves isolated the sample bulb. Mhen the sample was collected, the sample lines were purged.
Sample analysis was conducted by using a Carle Model 8700 gas chromatograph. Instrument specifications are presented in Table A-1. This unit was equipped with a thermal conductivity device. The output was recorded with a Linear Instruments Model 252 dual pin recorder. All samples and calibration standards were analyzed using repeat injections. A ten foot by 1/8 inch 0. Q. stainless steel column packed with Molecular Seive 13x, 80/100 mesh operating at 195'F, was used to separate the component gases. The gas chromatograph operating
. conditions are presented in Table A-.2. Calibration standards consisted of e several known concentrations of hydrogen (0.595%, 5.12K, 13.19K, 18.27K,
balance nitrogen) and oxygen (5.0X, 18.3X, balance nitrogen}. These standards were analyzed at the beginning and end of each sampling day.
TABLE A"1 GAS CHROMATOGRAPHY SPECIFICATIONS Sens i ti vi ty Variable, two level detector sensitivity switch Attenuator Eleven step binary type, 1024 to 1 Detectors Dual chamber, 100 1 volume with 8-10K, 0.013 dia. matched thermistors Power Requirements 115V, 60 Hz Temperature Control Ambient to 200'C
TABLE A-2 GAS CHROMATOGRAPHY OPERATING CONDITONS Argon Carrier Gas Flow 20 ml/mi.n 8 25 psi Oven Temperature 90 C Column 10'S tubing with molecular sieve 13x, 80/100 mesh Oetector Temperature Low Position Attenuation Variable
1 4 Attachment No. 4 to AEP:NRC:00500G Donald C. Cooing Nuclear Plant Unit, Nos. Additional Information on Hydrogen l and Mitigation and 2 Control Report on Igniter Performance Studies}}