TEE UNIVERSITY OF MICHIGAN INDUSTRY PROGRAM OF THE COLLEGE OF ENGINEERING A STUDY OF POOL BOILING IN AN ACCELERATING SYSTEM Herman Merte, Jr. A dissertation submitted in partial fulfillment of the requirements for the degree of Doctor of Philosophy in The University of Michigan 1959 December, 1959 IP-408

Doctoral Committee: Professor John Ao Clark, Chairman Associate Professor Herbert H' Alvord Professor Julius To Banchero Associate Professor Kenneth F. Gordon Professor Frank LB, Schwartz Professor Gordon J. Van Wylen ii

ACKNOWLEDGEMENT The writer desires to express his gratitude to Professor John A. Clark, Chairman of the Doctoral Committee, for giving so freely of his time with advice, encouragement and guidance throughout the course of this investigation. The writer also expresses his appreciation to Professor Gordon J. Van Wylen for his constant encouragement and advice, and to Professors Herbert H. Alvord, Julius T. Banchero, Kenneth F. Gordon, and Frank L. Schwartz for their interest and cooperation. The financial assistance of the United States Army Ballistic Missile Agency, Redstone Arsenal, Alabama,, the General Electric Fellowship Program, and the Horace H. Rackham Graduate Fellowships are gratefully acknowledged. The assistance of Mrs. Norlene Martin and the cooperation of the Industry Program of the College of Engineering in the preparttion of the manuscript is highly appreciated. iii

TABLE OF CONTENTS Page ACKNOWLEDGMENTb ~ *. Q * o o 4 4 *r 4 4 iii LIST OF TABLES 0 e a e a o0 0 a Q oooo o o a o a o o. 0 4 o o. 0 0 o a S * 0 Vi LIST OF FIGURES a a a a a a a 0a0 a Oe a a O a a a a a a Oa a o a a a a a a o0 a 0 a a O 0 4 Vii NOMENCLATUREaa aaa.ao oa.foaooa aa~OaO xi I INTRODUCTION a a a a4 0 a o. a 0 a a a a a a a a a a a O a a 0 a a a a aO O a a O a a a a a 1 A. Purpose a......ooo... ao.oo......oo0 ooooo.oo....o.o 1 B., The Boiling Phenomenaa a o a a. a o a.oa a. a 0a a a aO o0 a 0 o a a 1 B. The Boiling Pheno [[[ [[[ [[ [[ [ 1 C. Acceleration of a Boiling System....,,,,,...,,,...... 4 D. Existing Relationships in the Literature Showing an Effect of Acceleration. *.......*........oo.... 7 E. Previous Experimental Work in Boiling Heat Transfer Embodying an Alteration of the Gravitational Field.. 10 II. EXPERIMENTAL APPARATUS AND INSTRUMENTATION. O OO....O. oo. 12 A. Configuration. 40..00 00...00f.v* 00,o 0 00000000 0000 12 B. Construction0 0000 0040000 00 00 000000 0 12 C. Instrumentation.................................... 32 III. ESTIMATION OF ERRORS 0 34 A. Temperature..000.0........0..................o..... 34 B. Rotation of the Main Shaft.~.oo..o..o.,.oooa 49 C. Acceleration at the Heating Surface.........*0....... 49 D. Pressure on the Heating Surface.............o...**. 53 E. Heat Flux *.0. a0aa 0aaoaaoOOOO la*aa*oao 000 0a 00a0 57 F. Calculation of Heat Flux Rate from Temperature Gradient in Heater Block...0...0000*.000. 59 G. Measurement of Barometric Pressure......*..,.0* o 0. 62 H. Measurement of Specific Resistivity of Water in the Test Vessel,.... 0..0. 0 o 0. o. o o. 62 IV. TEST PROCEDURESo0o04 Oo 0 o000 00 o00 0 0 0 0 0 0 0 0 0 o 63 A. Fluid O 600000 00 B. Heat Flux Range Covered...O... 0. o......00.00o00...0o 64 C. Heater Surface Treatment0....*....0 0*....O... 64 D. Range of Accelerations Covered..0....... oo0.... o0... 65 E. Locatidn of Water Temperature Thermocoupleso..00.000.. 66 iv

TABLE OF CONTENTS (CONT'D) Page F. Variation of Cooling Water.,....0,.0....,.......o.... 66 Go Attainment of Steady State Conditions.., oe....,....o- 68 H.o Criteria for Acceptable Data......................... 69 V. TEST RESULTS,0..O..O........ OOO......OO...e....o........ 70 A0 Natural Convection.,0000*,............ 70 B. Boiling.o.........e... oo....o.oooo.............oooo..o 74 C. Overall Results..., 101 VI. ANALYSISO. o O * * * 6 0 * * 117 A. General.. *a....0 0..0. 00. a.......... 00 0.. 117 B. Bubble Relationships in Boiling0.o.......0.......... 118 C. Some Observations on Boiling at Large Values of Heat Flux......... 0 0 e 0 0 0 0 0 @ 0. 123 D. The Influence of Acceleration on Boiling Area........ 129 E. The Influence of Acceleration on the Number of Nucleating Sites..............*...*....... 138 F, Concluding Remarks*.......,.O...................... 142 APPENDICES A, DERIVATION OF EQUATION FOR ERROR IN MEASURING WATER TEMPERATURES... oo..................... a. o060 143 B. DERIVATION OF EQUATION FOR HEAT LOSS BY CONDUCTION THROUGH HEATER SKIRT.................o... 147 C. APPROXIMATION OF VELOCITY OF FLUID AT HEATING SURFACE DUE TO CORIOLIS' FORCE. O....O *4....a. *.OO e...O.oo 153 DO DATA..0*006000000*6 155 BIBLIOGRAPHY.. oo. o.,s. o. o..o.0.o..e0 0..eoee o o oo.o...eee167

LIST OF TABLES Table Page I List of Mercury Thermocouple Slip-Ring Channels and Radius from Centerline of Rotation................ 36 II List of Mercury Channel Pairs Associated with Each Thermocouple Circuit..,,.....,........... 37 III Values of Uncertainty W(Tw -T1) Calculated from Typical Test Data aa........I...n*.. 0....*.....0....... 44 IV Values of Parameters in Test Vessel Which Vary with Rotational Speed..................... 52 V Loss by Conduction Through Skirt for Convection Tests..... 61 VI Sample Sequence of Accelerations During Test Runs to Determine Effect of Past History of Acceleration.......... 66 VII Comparison of Heat Flux Calculated from Heater Temperature Gradient with that Calculated from Power Measurement. 111 vi

LIST OF FIGURES Figure Page 1 Orientation of Acceleration with Respect to Heating Surface........................ c *...... 13 2 Test Vessel,.................... 14 3 Centrifuge Assembly..,. a 15 4 Detail of Copper Heater-Skirt Intersection,............. 17 5 Heater Thermocouple Locations.........,.,...... 18 6 Underside View of Heater Block..,.,..........*.,, 20 7 Cover Installed on Heater Block........................... 20 8 Guard Heater Assembly,.............................. 21 9 Assembled Heater Unit............... b........... 21 10 Assembly of Inner Container Walls and Heater............... 22 11 Inner Container Test Assembly..................... 22 12 Test Vessel Installed in Centrifuge.,...,,,o-..... 24 13 Construction of Heater Thermocouple Probe.............. 24 14 Equivalent Thermocouple Circuits........................* 27 15 Sketch of Mercury Thermocouple Slip-Ring Assembly.......... 28 16 View of Mercury Thermocouple Slip-Ring Assembly........,,,,,,, 31 17 Overall View of Test Apparatus...................... 31 18 Typical Plot Indicating Effect of Rotation Upon Thermocouple Readings..,............*....'.a........ 38 19 Corrections Applied to Thermocouples to Compensate for Rotation o Heat@~e Trcpsle4* * O * 40 20 Location of Heater Thermocouple Holes.....,,............... 42 vii

LIST OF FIGURES (CONT'D) Figure Page 21 Typical Conditions for Thermocouple T with System Subjected to Acceleration 21,15 Times normal Gravity,....... 46 22 Test Vessel Dimensions for Calculation of Acceleration at the Heating Surface..............o..............e 51 23 Relation of Heater and Liquid Surfaces Due to Rotation...... 55 24 Estimated Heat Loss by Conduction Through the Heater Skirt for Various Total Heat Fluxes................... 60 25 Various Locations of Thermocouple T6 for Different Test Runs *..~...@s.........................4...... 67 26 Plot of ATc Versus Acceleration for Natural Convection Indicating the Effect of the Flow Guide......... 71 27 Run No. C-5, Temperature Data for Convection as Taken...... 72 28 Correlation of Natural Convection with Acceleration Normal to Heating Surface... *.........***.....*........ *..... 73 29 Run No. B-15. Plot of Tw-Tsat and Tw-T5 Versus Water Temperature T.*.......................*...............*. 75 30 Run No. B-15. Plot of Tw-Tsat and AT in Heater Block Versus Acceleration with Pool Boiling..e.O.*.o. 82 31 Run No. B-15. Difference in Water Temperatures for Various Accelerations and Subcooling8...3......... 83 32 Run No. B-15. Temperature Profile Between Heater and Water Surface for Various Accelerations..................... 84 33 Run No. B-9. Tw-Tsat and Tw-T5 vs. Water Temperature T5 at Various Accelerations........................... 85 34 Run No. B-9!. Temperature Profile Between Heater and Water Surface for Various Accelerations.................... 89 35 Run No. B-9. Difference in Water Temperature for Various Accelerations and Subcooling................ 91 36 Run No. B-9. Plot of Tw-Tsat and AT in Heater Block vs. Acceleration with Pool Boiling............ 92 viii

LIST OF FIGURES (CONT ID) Figure Page 37 Run No, B-14. Plot of T -T at and Water Temperatures vs. Acceleration with Pool or9lingo,,o,,,,.......... 94 38 Run No. B-14. Plot of AT in Heater Block Vs. Acceleration., 95 39 Run No. B-14. Temperature Profile Between Heater and Water Surface for Various Accelerations,.,,.......,.. v.,,.,. 96 40 Run No. B-22. Plot of T -T and AT in Heater Block vs. w $at Acceleration with Pool Boilng.....G..................... 98 41 Plot of Water Temperature vs, Acceleration for Two Runs Identical Except for Location of Thermocouple T6.......... 99 42 Run No. B-22, Temperature Profile Between Heater and Water Surface for Various Accelerations.,,..,.........,......... 100 43 Run No. B-21. Plot of Tw-Tsat and Water Temperature vs. Acceleration with Pool Boiling,...o,,......,..,.Ac,,,,... o 102 44 Run No. B-21, Plot of AZT in Heater Block vs. Acceleration.. 103 45 Run No. B-21, Temperature Profile Between Heater and Water Surface for Various Accelerations......a o.B e e o. v.... 104 46 Plot of q/A vs. Tw-Tsat for Boiling in Standard Gravitational Field....aa..a......*...*....................* lo 106 47 Influence of Acceleration on Tw-Tsat with Pool Boiling to Saturated Water....*..........,.....o................. 107 48 Plot of Tw-T5 vs. Acceleration with Pool Boiling to Saaturated WaTerd oo.................O.......o............ 109og 49 Subcooling at Thermocouple T5 as a Function of Heat Flux and Acceleration,.,,..............., *........, v.... v. 113 50 Example of Coriolis Acceleration on a Particle Constrained to Move Radially on a Rotating Bar,,,,,,,,,. o,,,,,,.... 113 51 Correlation of Natural Convection with Flow Guide...,..... 116 52 Representation of Hirano and Nishikawa(49) for Boiling Heat Transfer........,.................. o, 119 ix

LIST OF FIGURES (CONT D) Figure Page 53 Interdependence and Complexity of Boiling Elements...,...... 119 54 Thickness of Boundary Layer on a Ho izontal Heating Surface as Measured by Hirano and Nishikawa 54) Using Refraction Method' A a.,...... * 122 55 Effect of Heat Flux on Fluid Temperature Near Heating Surface vith Forced Convection Boiling, Due to Treschov,55)..* O &....,.*.WW.OW... 122 56 Representative Plots of q/A vs. AZ Near Peak Heat Flux...... 125 57 Illustration of Area of Influence of Bubble with Boiling Heat Transfer..,................................. 131 58 Convective Fluid Flow Pattern in the Experimental System with and without Flow Guide..,.....,.,....,... 131 59 Effect of Acceleration on Peak Heat Flux with Pool Boiling.. 135 60o Calculated Values of y as a Function of Total Heat Flux and Acceleration.................................... 136 61 Cross Plot of Figure 601........7........................ 137 62 View of Heater Surface After Test Run with Slightly Contaminated Water at Flux q/A = 99,500 Btu/hr-ft..... 139 63 Number of Nucleating Sites as a Function of Heat Flux and Acceleration.......................................... 141 64 Schematic of Thermocouple Tube Shown in Figure 21........ 144 65 Extended Surface Representing the Heater Skirt............... 148 66 Model Used to Calculate Maximum Possible Water Velocity Due to Coriolis Acceleration in Test Vessel........1........ 154

NOMENCLATURE Other nomenclature is defined locally as necessary a Acceleration normal to heating surface a/g Dimensionless acceleration A Area C Constants Cp Specific heat D Diameter DNB Departure from Nucleate Boiling, defined in Chapter VI-C f Frequency of bubble formation g Local graVitational acceleration gc Mass-force conversion constant = 32.174 lbm/lbf ft/sec& h Convective heat transfer coefficient Gr Grashof number Nu Nusselt number Pr Prandtl number N/A Active nucleating sites per unit area q/A Heat flux rate T Temperature Tn Temperature at thermocouple n AT(m-n) Temperature difference = Tt - Th W Uncertainty Thermal diffus ivity Volumetric coefficient of thermal expansion xi

Y Defined by Equation 66 6 Boundary layer thickness CY Surface tension X Latent heat of vaporization Defined by Equation 63 AG Tw - Tsat Contact angle Absolute viscosity v Kinematic viscosity Subscripts b Bubble c Convection f Film 1 Liquid Sat. Saturation t Total v Vapor w Wall xii

I. INTRODUCTION Ao Purpose The purpose of this study is to investigate the effect of acceleration on a system in which boiling is taking.place from a flat heated surface. Acceleration of the system provides the equivalent to a change in the force field acting on the system. An understanding.of the role of the force field in the boiling heat transfer process may add to the overall understanding of the boiling phenomena and increase its effective application. B, The Boiling Phenomena The characteristics of the three regimes of boiling are abundantly available in the literature The maximum heat transfer rate attainable with nucleate boiling is designated as the peak beat flux or "burnout point", and has been characterized as the condition where the bubble population on the heating surface is so great that (4-6) the bubbles interfer with one another( A further increase in temperature difference between the liquid and the heating surface then results in a decrease of heat transfer rate because of this interference, until a minimum point is reached, after which stable film boiling is present. The higher heat transfer rates associated with nucleate boiling over that of convection have been ascribed to the action of the bubbles as agitators in the "laminar sub-layer" rather than as media of energy transport 7) The wide variation of temperature

difference observed on an electrically heated surface with patchwise boiling would tend to confirm this. The mechanism of nucleate boiling has been considered in the following stages: 1. The presence or formation of a nucleus from which growth of a bubble can take place. It has been shown that because of surface tension effects (2) a minimum value of superheat in the liquid is necessary in order for a bubble nucleus of finite dimensions to form and grow. For a given superheat, the minimum size of the nucleus which is unstable and hence will grow has been termed the critical size.(l) The smaller the surface tension, the lower is the superheat required for the formation of such a nucleus of given size (i.e. with a lower superheat the nuclei are more apt to form for a given superheat). The roughness of the heating surface also affects the ease with which nuclei may be formed.(9) At lower pressures surface boiling is initiated at a lower surface temperature with the presence of a dissolved gas (27), which provides a source of nuclei. 2. The early and mid growth period of the bubble. Consideration of the hydrodynamic, heat transfer and surface tension effects on the growing nucleus(ll 13) has indicated that owing to the surface tension the initial growth rate of the bubble from the critical size is small. Once the bubble has reached a size such that the surface tension effect is no longer important, the growth rate becomes very large. This is roughly that size which is discernable

-3 -to the eye. In this stage the rate of growth is governed by the rate of heat transfer across a thin liquid film surrounding the bubble, and hence is dependent upon the degree of superheat of the liquid. 3. The late growth and departure period of the bubble. In this period the growth of the bubble has "consumed" the liquid superheat in the immediate vicinity of the bubble and the bubble ceases growing. Depending upon whether the bulk liquid is subcooled or saturated, the bubble may then either collapse immediately, depart from the heating surface and then collapse, or depart and rise, due to dynamic and buoyant forces. It has been demonstrated in nucleate boiling that bubbles form at preferential active points on the heating surface.(2,9) These active points are postulated to be due to minute cavities existing in the surface which trap vapor from an earlier bubble which has departed(9, or simply the remnants of an earlier bubble(14), which then serve as nuclei for further growth. A theoretical study of nucleation(21) in the presence of normal gravity has lead to the conclusion that nucleation always occurs at boundaries of gas or vapor trapped in surface (38) cavities. A microscopic photographic study of boiling confirmed this. Experiments in which the effect of the entrapped vapor had been minimized or eliminated enabled extremely high values of superheat to be (22-2[) () attained. However, it has also been theoretically demonstrated (43) that even if no cavities exist on the heating surface, nucleation will take place preferentially at the surface if any finite degree of

-4 -wettability exists between the fluid and the surface. Increasing the surface temperature results in more active nuclei and an increase in agitation, and hence an increase in heat flux until the peak heat flux is reached. The literature is not clear as to whether the increase in agitation results solely from an increase in the number of active nuclei or if each existing nucleus also contributes more agitation because of the possible greater liquid superheat in the vicinity of the heating surface. It is likely both factors contribute to the effect. C. Acceleration of a Boiling System Consider the boiling system represented in Figure 1 with a spherical vapor bubble just attached to the heating surface. If the system is accelerated as shown in a direction normal to the heating surface in a standard gravitational field it can be shown that the net buoyant force acting on the bubble will be: a FB = Vb(PQ - Pv) g + PVb (1) At the instant the bubble is detached from the surface it can further be shown that the acceleration of the bubble with respect to the heating surface will be: aB = Pv (g + a) (2) Pv

If the effect of surface tension is the only force holding the bubble to the surface it is obvious from Equation (1) that a smaller bubble volume will suffice to overcome a given value of the adhering force with the system under acceleration. This effect is inherent in (14) the following expression developed by Fritz for the maximum volume of a bubble at departure from a surface: 3/2 2l v Vb(max)= (0 0119 )3 g, (3) Equation (3), based on the work of Bashforth and Adams(25) and Wark(26) considers the equilibrium of the surface of curvature separating 2 phases in a normal force field of standard gravitational acceleration, Assuming a constant contact angle, an increase in the gravitational acceleration will decrease the maximum volume of the bubble at departure from a heated surface. If the effect of an imposed normal acceleration of the boiling system were to cause the bubbles to be detached prematurely in the sense of Equation (3), then a number of postulations of a general nature might be made at this point, With an electrically heated surface the time averaged heat flux must remain constant, and any decrease in agitation due to the smaller bubble sizes must be compensated for in other ways, The frequency of formation of the bubbles at each active point may increase. If this is insufficient, the number of active sites may increase, which most likely would require an increase in the surface temperature, If, on the other hand, the major agitation caused by the bubbles bakes place during the early growth period, smaller bubble sizes will either

have little effect if the frequency of bubble formation remains unchanged or will act to decrease the time averaged surface temperature if the frequency increases. A complicating feature of the consequence of smaller bubble sizes will be the increasing contribution of natural convection, both because of the increased area available and because of the increased force field resulting from the imposed acceleration. The peak heat flux might also be expected to increase under the action of the acceleration. The agency causing the departure of the bubble from the heater surface will probably be one of the major factors influencing this effect. The effect of buoyancy due to normal gravity has been discounted as an explanation for the departure of a vapor bubble from a heated surface since it has been observed that vapor bubbles may be ejected even from the lower side of a horizontal surface. A plausible cause of departure is (18,60) given which attributes the motion to the inertia of the surrounding liquid. During the rapid growth period of the bubble the surrounding liquid acquires momentum. As the growth rate decreases because of decrease in superheat, the inertia of the liquid causes a reduced pressure field on the upper surface of the bubble tending to draw it from the surface. Photographs upon which this description is in part based showed that during the greater portion of growth the bubble maintains a hemispherical shape, but just prior to departure it becomes elongated perpendicular to the heating surface. The effect of surface tension in nucleate boiling, discussed below affirms this view. It has been observed that the introduction of small quantities of additives which decreased the surface tension decreased the peak heat

(87 -flux and, for a given temperature difference at less than peak heat flux, increased the heat flux 19,20) The latter effect seems logical in view of the effect of surface tension on the tendency for nucleation. Such a condition requires less liquid superheat to cause nucleation and hence results in lower bubble growth rates, If the liquid inertia is important in drawing the bubble from the surface then lower growth rates mean slower departures, which in turn mean a greater tendency for the heating surface to become vapor bound, Hence, the peak heat flux would have a lower value, as was observed. A discussion similar to the above is given by Larson 28) A recent photographic study of boiling in the absence of gravity, however, seems to invalidate the importance of liquid inertia as a mechanism in detaching the bubble from the heating surface. The vapor remained adjacent to the heating surface, and there was no evidence of bubbles being pushed away from the surface to any appreciable extent during their formation. D. Existing.Relationships in the Literature Showing an Effect of Acceleration 1, Several. correlations for boiling heat transfer which incorporate a term for the gravitational acceleration exist in the literature~ The relationship due to Rohsenow(29)~ Cpz_ C __ g 1a 1.7 Csf (4) X fIAIkX g(PYPv) r contains an acceleration term as a result of using Fritzv relation for the bubble diameter in a bubble Reynold's Number, Rewriting Equation (4) we

-8 -have q/A = I 5HY- Z l e L ]1 (5) In extending this correlation to a boiling system undergoing an acceleration, its validity will depend upon the continuing fidelity of the assumptions made. Among these assumptions are: (a) The effect of the contact angle cp, which is included in the constant Csf. (b) The relationship between the frequency of bubble formation and the diameter as it leaves an active site, f - Db = constant. (c) The overall heat transfer is proportional to the heat transfer to the bubbles while attached to the surface. 2, Gravitational acceleration appears in the expression for boiling heat transfer derived by Chang(3Q0) given in Equation (6), as a result of the extension of the wave analysis of natural convection to boiling. C1 and n are specified as constants to be determined by experiment hb.16 k[l + Pr {C1(q A X p 1 i j L Pr 6gp2 This equation is not explicit in q/A, but if BrC'/q/A. v n (? P W %: n> (1 -P) (7) by at least an order of magnitude, it can be written with small error as q/A = K1 g1/3-2n (a )4/3-s2 (8)

where K1 is a function of properties given in one form by K o56 clp(4) l/ / C (9) In attempting to utilize Equation (8) to describe the boiling characteristics in an accelerating system both C1 and n, and hence K1 may be a function of the acceleration. 3o A number of relationships for the peak heat flux with pool boiling have been presented in the literature~ (a) Addoms, whose work is cited in McAdams (lpp 384) correlated experimental peak heat flux data by plotting (q/A)- o P versus - P Pv(gU,)l/3 Pv If the deviation from a straight line near the critical state is ignored, the best line of the data yields an expression of the form (q/A)p = C X Pv (ga%)l/3( )n (10) Pv where C = 2 and n = 1/2 (b) By the application of dimensional analysis Borishanskii derived similarity criteria for the condition of peak heat flux which, in conjunction with experimental data, result in the following: (q/A)p = K2 ( ) 1/2 [ g(p P.. ]v/4 (11) where K2 = 0o13 + 4N~4 (12)

-10 -and N = 2 (13) '[g(p - Pv) ] N is a correction factor to K2. For water boiling at atmospheric pressure the term 4N~4 has a value of approximately.013, or 10% of 0.13. (c) By considering the peak heat flux as a hydrodynamic (5) stability problem, Zuber arrived at Equation (14) without the necessity of an empirical constant 1/4 p + pI 1/2 (q/A)p - 2 X (Pv)/ [~ g(Pg - Pv)] [ 1/2 (14) (d) Using two different models for the peak heat flux Chang and Snyder(44) obtained correlations identical to Equation (11) with K2 = 0.145. (e) The correlation for peak heat flux for both pool boiling and forced convection derived by Griffith(45) specifies that the peak heat flux is proportional to gravitational acceleration with an exponent of 1/3. It is noted that except for (a) and (e) above, the peak heat flux is given as proportional to gravitational acceleration with an exponent of 1/4. E. Previous Experimental Work in Boiling Heat Transfer Embodying an Alteration of the Gravitational Field (32) In a paper referred to earlier, Siegel and Usiskin made a photographic study of boiling from several heater configurations in the absence of a gravitational field. No attempts were made to measure heat fluxes or temperatures for the series of tests reported. With water flowing in a vortex in an electrically heated tube(33) a peak heat flux on the order of 55 x 106 Btu/hr.-ft2 was attained.

-11 -This was attributed to the effect of the centrifugal acceleration on the bubbles forming at the heating surface, estimated to be 18,000 times normal gravity at the exit from the heating tube. However, the contribution of forced convection could not be isolated, A number of investigations have been made on the effectiveness of an increased force field in promoting heat transfer by natural convection and condensation.,

II, EXPERIMENTAL APPARATUS AND INSTRUMENTATION The experimental work was performed using the facilities of the Heat Transfer and Thermodynamics Laboratory in the Mechanical Engineering Department. A. Configuration In order to most effectively isolate the influence of the increase in buoyant force on the bubbles with nucleate boiling in an accelerating system, the orientation of a heater surface as shown in Figure 1 was selected for the experimental work, that is, with the acceleration applied normal to the surface. For practical reasons the centrifuge principle was used to obtain the acceleration. To maintain an approximately uniform depth of liquid over the surface and the acceleration normal to the surface at all speeds it was necessary that the vessel be pivoted, as in a fly-ball governor. Figure 2 is a drawing of the pivoted test vessel and Figure 3 shows the overall centrifuge assembly. If the center of gravity of the test vessel is located at the surface of the liquid, a plane tangent to the liquid sunface at the center of the vessel will be parallel to the heater surface. The liquid surface will have the shape of a parabola of revolution. If the center of gravity of the test vessel is located at the heater surface the acceleration will be normal to it. Obviously both conditions cannot be maintained for a finite depth of liquid, and a compromise is necessary. This aspect of the apparatus is discussed in more detail in Chapter III. B. Construction 1. Heater Assembly The heater itself consists of a cylindrical piece of copper, containing 1% lead for machineability, 3 inches in diameter and 1 inch

-13 --~ a 9 HEATING SURFACE Figure 1. Orientation of Acceleration with Respect to Heating Surface.

-14 -PIVOT ARM TC -6, WATER /TC- 5, WATER CONDENSER COI L OUTER CONTAINER DRIP PLATE "' 08~a~ vt-3~ XFLOW GUIDE [ " J~ I P1 T l ~ [ ~ ~ INNER CONTAINER I I 11 I t < TC-I, HEATER HEATER BLOCK ~_ | - - CHROMEL RESISTANCE RIBBON 94 - / /o W d / HEATER COVER:- s~3 \ \\\l:' FIBERGLAS ep 1;i' ' ~ ~ INSULATION.7 ~?., ~G~ -9 ~~~~~S.S. SUPPORT p d o| Blc" S.S. SPACER RING |'kf M M M \L u M u LI LI EAT i HEATER GUARD cb,tgGUARD HEATER 43, _ L ELEMENT \T-G, DIFFERENTIAL THERMOCOUPLE Figure 2. Test Vessel.

COOLING WATER INLET TAC HOM#ETER GENERATOR (TO ELECTRONIC COUNTER) BELT DRIVE KEROSENE THERMOCOUPLE COMPENSATION BATH MERCURY THERMOCOUPLE SLIP-RING ASS'Y 10 CHANNEL S 11www w WI VARIABLE - SPEED HYDRAULIC TRANSMISSION COUNTERWEIGHT i MAI N SHAFT~ COOLING "'*'\ \ELECTRIC WATER OUTLET DRIVE MOTOR TEST VESSEL POWER SLIP-RING ASSry r0DRAIN Figure 3 Centrifuge Assembly.

-16 -long. One end of the cylinder serves as the heat transfer surface. In order to provide a continuous surface and to keep the heat loss by conduction to minimum, this end was undercut with a bevel and a mating piece machined from stainless steel.002 inch undersize, as shown in Figure 4. By immersing the copper in liquid nitrogen the stainless steel skirt slipped over the bevel, and upon return to normal temperature a water tight smooth surface was obtained. The heating surface was then chromium plated. Four 1/32 inch diameter holes were drilled radially into the cylinder for the insertion of thermocouple probes. Two of the holes extended to the center from opposite sides as shown in Figure 2, the centerline of one being approximately 1/16 inches below the heater surface and the other being 7/16 inches below the heater surface. The two remaining holes were located at 90~ from the above holes, extended half-way to the centerline of the cylinder, and were also located 1/16 inches and 7/16 inches below the heating surface, Figure 5 shows the location of these holes with the respective thermocouple designations. In the lower end of the cylinder 32 parallel slots 0.008 inches wide by 5/16 inches deep were machined to accomodate 6 feet of Chromel A heater ribbon 1/4 inch wide by 0.002 inches thick. The ribbon was insulated from the copper with 0.003 inch thick strips of mica. Gold foil 0.001 inch thick shunted the ribbon where it emerged from one slot and entered the next in order to minimize "hot spots." Figure 6 is a view of the underside of the heater showing the slots and one of the thermocouple holes near the stainless steel skirt. To protect the ribbon

-17 -HEATER SURFACE 026S.S. SKIRT 2.946",iA Fgr1 De1 /16" Figure 4. Detail of Copper Heater-Skirt Intersection.

3" DIA TC-I (NEAR SIDE) TC-3(NEAR SIDE).- 7 TC-4(FAR SIDE) TC-2(FAR SIDE) Figure 5. Heater Thermocouple Locations.

and to provide a uniform temperature on the underside of the heater a nickel and chromium plated copper cover 1/4 inch in thickness was assembled to the heater block as shown in Figure 7, In order to minimize heat losses by radiation and convection from the underside, the heater is installed in a 1/4 inch thick chromium plated copper base, seen in Figure 8, which serves as a guard heater0 The thin stainless steel spacer ring (refer to Figure 2) supports the underside.of the heater cover. Differential thermocouples are installed between the heater cover and the heater guard (Figure 2), and a heating element located under it is. controlled to keep the temperature difference at a minimum. The assembled heater unit, Figure 9, is quite compact, measuring 6.1/2 inches in diameter and 2 inches in depth0 2, Inner Container and Cover The inner container side walls consist of a double wall sheet stainless steel welded assembly separated by an air space0 The inner surface was chromium plated to prevent contamination of the water due to corrosion at the weldsO Figure 10 is a view of the inner container walls attached to the heater assembly, Fittings for installation of the water temperature thermocouples are installed in the cover, and a coil of 3/8 inch O.D. copper tubing is silver-soldered to the underside as a condenser. This assembly was also chromium plated, Teflon gaskets are used wherever contact with the test water is necessary, The inner chamber is vented to the atmosphere thru a small tube which is surrounded by the cooling water as it enters the condenser. To prevent a change in pressure within the vessel due to an aspiration effect upon rotation, the

-20 -''s::':,,,',:i:::',,','...... _ i\ Figure 7. Cover Installed on Heater Block MINI:~:~:j:~::i:~ii-i:..:::.:::::::::.:::.:::''':.'~~i:i:~:ii: ON~ ~~ ~~~~~~~~~~~~~~~~~~~~:::::::,:::iii.Ii~~:::::::::i:::::::::iil:i:i ~ ii~iii~ii~::~~~::iii~~~:iii

-21 -Figure 8. Guard Heater Assembly. Figure 9. Assembled Heater Unit.

-22 -Figure 10. Assembly of Inner Container Walls and Heater Figure 11. Inner Container Test Assembly.

-23 -vent tube is connected by rubber tubing to a point near the centerline of rotation. Figure 11 shows the entire test vessel assembly with thermocouple and heater leads prior to installation in the welded aluminum outer container shown in.Figure 12, 3. Flow Guide The heating surface consists of a heated.section in the center of an unheated section, In order to eliminate the possible influence of the gross convection currents on the bubbles forming near the edge of the heated section, a flow guide.was constructed, consisting of a cylinder of thin sheet stainless steel open at both ends, The inner diameter of the flow guide is about 1/32 inch larger in diameter than the heated section, and is installed so that it rests over the heater on the skirta The net effect is that of having a test vessel whose entire bottom surface is heated, Several test runs were made with this flow guide removed. 4. Thermocouple Construction A representative thermocouple installed in the 1/32 inch diameter holes in the heater block is shown in Figure 13, These were constructed by filling a short piece of 1/32 inch OD, copper tubing with a ceramic cement, passing a 30 gage constantan wire thru the tube, and welding the tube and wire near the tip by pinching the tube, after applying a voltage with a charged condenser. The tip was then trimmed so that the junction occurred at the extreme end of the tube, A 30 gage copper wire was welded to the other end of the tube on the outer side in a similar fashion,

-24 -Figure 12. Test Vessel Installed in Centrifuge. 30 GAGE COPPER WIRE WELDED JUNCTION WELDED TO TUBE 1/32"0. D. COPPER TUBE /SPACE FILLED WITH Figure 13. Construction of Heater Thermocouple Probe.

The water temperature probe designated T5 in Figure 2 consLsts of a 30 gage copper constantan thermocouple cemented inside a 1/16 inch OD. stainless steel tube closed at one end and extending 1 Inch frocm a ' /8 inch OoD, s.s. tube for rigidity0 The thermocouple tip is located. 1/4 inches from the heater surface in the center0, The small extension tube is used to keep errors due to conduction in the tube to a miYnimmo., Water temperature thermocouple T6 is similarly constructed except that only 1/8 inch O,Do s.s. tubing is used for the outer casing. Several were made with different shapes to permit water temperatu:re measurements at.different locations~ 5o Thermocouple circuit In order to remove the thermocouple EMF's from the rotating member, a mercury slip.ring was selected as being capable of the greatest precision, A similar assembly was used by Fultz.and Nakagawa(39) with copper-cQns'tantan thermocouples, It was stated the spurious EMFIs gave no trouble, even with voltages measured on the order of several microvolts, The temperature errors in the present work are discussed in Chapter IISo It was found that 30 gage copper wires moving through 4the mercury deteriorated completely due to amalgamation after about 1/2 hour of operation~ Iron wire was the most commonly available material which would not be attacked by mercury, and 24 gage wire was introduced as an intermediary between the copper and mercury both on the moving and stationary sections. By the thermoelectric "law of intermediate metals, no error should result if the corresponding junctions on the rotating and stationary members are maintained at the same temperat;lire0

-26 -The equivalent circuit for the measurement of temperature with respect to the ice point is shown in Figure 14a. Ten concentric mercury channels were provided in a plexiglas piece, shown schematically with two channels in Figure 15. The stationary iron wires enter the mercury through the bottom, and the moving wires dip into the mercury through small holes in a plexiglas dust cover which rotates with the shaft. The corresponding iron-mercury junctions are the stationary and moving wires in the same channel of mercury. The mercury is maintained at a uniform temperature by having the plexiglas assembly rest on a heavy block of aluminum, which also serves as the upper bearing block, and by the stirring action of the wire as itn-rves through the mercury. A difference in temperature between the moving and stationary iron wires might be anticipated because of the stagnation effect, but calculations showed it to be negligible at the rotational speeds employed. The corresponding copper-iron junctions are maintained at a uniform temperature by insertion in a circular kerosene compensation bath which rests on the dust cover and rotates with it. Again, the stirring action of the thermocouples is relied upon to provide the uniform temperature. The water thermocouples T5 and T6, and the heater thermocouple T1 located in the center of the block near the heating surface (Figure 5) are used with respect to the ice point, and employ a common constantan wire in passing through the mercury slip ring. The remaining thermocouples in the heater block (Figure 5) are measured as differences as follows, by

-27 -Junctions in "Isothermal" Bath Copper g Fe Copper To \ Measuring Potentiome ter Junc ti on Copper I Const Fe I Hg Fe Const. Sttionary Rotating Section Section Reference Junction a. Equivalent Circuit for Measurement with Respect to Ice Point. Copper, Fe I Hg t Fe 1 Const. FTo Temperature Potentiometer Lopper Difference to Copper A, Fe 1 Hg I Fe 1 Const. be Measured Stationary Section Rotating Section b. Equivalent Circuit for Measurement of Differences. Figure 14. Equivalent Thermocouple Circuits.

-28 -STATIONARY WIRE ROTATING WIRE MAIN SHAFT KEROSENE COMPENSATION BATH DUST COVER MERCURY (ONLY 2 OF 10 CHANNELS SHOWN) ALUMINUM BASE Figure 15. Sketch of Mercury Thermocouple Slip-Ring Assembly.

-29 -a suitable switching arrangement: T2 - T1 = T (2-1) T1 - T3 = AT (1-3) T2 - T4 = AT (2-4) Owing to the contact between the copper tubes serving as one of the thermocouple leads and the heater block it was necessary to use the circuit represented in Figure 14b to measure the voltage differences. In addition to the above thermocouple circuits, the differential thermocouples between the heater guard and the underside of the heater block are taken through the mercury channels. Also, to check for the effect of differences in velocity of the wires moving through the mercury, the innermost channel (radius = lol/8 inches and the outermost channel (radius = 3 15/16 inches) are short-circuited on the rotating assembly, and periodic measurements taken at the stationary wires at various RPM. All connections between various parts of the circuit are made in tight copper boxes to avoid air currents and other thermal gradients. Figure 16 is a view of the thermocouple slip-ring assembly showing the mercury channels, dust cover, kerosene compensation bath and support for the stationary junctions in the kerosene bath. The wires from the test assembly to the slip rings are led around the bearing through slots in the main shaft. 6. Cooling Water System Cooling water for the condenser is transmitted to the rotating system through a rotating seal at the upper end of the main

-30 -shaft. A hole drilled at the center of the shaft extends below the upper bearing, and connection with the pivoted test vessel is made with rubber tubing, as seen in Figure -12. The water then enters the lower part of the shaft, again via rubber tubing, bypasses the lower bearing through the hole drilled in the shaft, and enters a drain at the lower end of the shaft. For certain tests with low flux rates it was found necessary to drastically reduce and control accurately the rate of cooling water flow. For these tests small sharp-edge orifices were installed at the discharge from the lower end of the shaft and calibrated with a needle valve and pressure gage located near the stationary water inlet to the shaft. 7. Electric Power System Electric power for the main and guard heaters is transmitted to the rotating shaft via a slip ring assembly located under the lower bearing, as seen in Figure 3, The wires bypass the lower bearing through slots in the shaft and coiling of the wires prevents their imposing any restraint on the pivoted test section. 8, Power Drive The main shaft is rotated by means of a belt connected to a variable-speed hydraulic transmission, driven in turn by an electric motor. 9. Counterweight Both the test assembly and the counterweight are pivoted on a cross-arm attached to the main shaft. The counterweight consists of a

-31J.......a............ _ --..............................................S* a _ ~~~~~~~~~~~~~~~~~~~~~~~~~~.>; i ' S;d':~'E CRR-#-: 3DLRE;0 0:;.-dER- X E EEEE.:..........................................................# - 0.......:.0........S.... i:...... *a Ei i;:E:; i: ilil:i':;: g' e: E- gi ggE~i~#::::iitg::;;i '%l8sssssgi-.-i:::Ek8~8i ~ g:.::idl a8i i:;i:;i::ig-# #lliiiiiiii.: i 01 l. | a i i:.:-:i::E;:i. 1 _ a l ~~~~ ~~~~~~~~I::i-i liii iiii:'i -:.iiiii_ il iii fi i i:::g.i#:.: N... " '= i ~:S E i..............:. '......................:....::::.:: a........... F i g u r e.N. ii,.. -',',.'l.','.....EE....l i p.................:"' -.B~~~~~~~~~~~~~~~~~~~~' - g a::g:gg: -: fEEi-~ii: ii~:~:::E~SEi::EE EE'. R _ESER:'SE' iS E _ i _': E:'_ M;X.:-:-............................_ -# 1! _:.'-........... ab lity _E$0044ii$E:__=, - ti ii~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~i 9: i 0 _B B B B X iiiiiii ~~ii~iI::::::~::~:::: ~ {a B '; { _ E - ig::::'5 ig:E;S'i Fa,, - - i 200E - -E:: - > g g -- lf~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~iii:iiii::::::::~:::::::ii:::: '. '.iii '.B'..................................................:' _............,.:-.'......... ii..'-a,.... { i i; ' Si ia::i; l:;0: ii ii i~~~~~~~~~~~~~~~~~~~~~~~~~~~~i::tii | i::E';ii:; i: iii i: E: i: i ' 'g~> a a~ ~~~~~~~~~~~~i~~i~:: I~~lii:: iii iiiii il. i; i E B igg {,# ~.6~R B.; E i; i ii:gg;::; g g i:. i~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~li~~~~l.;.i-.E. iB Bii lii~iii~.1 X..........: E g E:; E::;: i R B _ _ _ 1 1 1................E E iE; E: E E i: __~~~~~~~~~~~~~~~~~~~~~~..,;...........:g 'f:L'..............ER'... a R.:...... '.''...... _~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~::: ""'':': ':R'''':'':':''''';:' _>il9|i-l~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~liiir~ Figure 16. View of Mercury Thermocouple SlipRing As sembly. iliilii~~~~~~~ii'~~~~:i~~~~i~~i~~~i~~i:::::::,:~::::::::::::::~:-i ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~:i'.,::,..''i. <.Ct,,,k.Y,. S.,:~:::~- i:''' Figure~~~~~~~~~~~~~~s _cacew Oveal Vie ji: Tes Apar'

threaded rod on which the desired weights are attached, with a pivot at one end. The counterweight was made equal in weight to the filled test assembly, and the center of gravity was made equal by a trial and error process, by moving the weight up and down the threaded rod until no vibration could be detected, Hence the system was dynamically balanced for all rotational speeds up to 220 PRM, the maximum for which it was designed. C. Instrumentation 1. Temperature All thermocouple EMF were measured with a Leeds & Northrup Type K-3 potentiometer with the exception of the guard heater differential thermocouple. It was considered adequate to maintain the temperature difference between the heater and guard within 5~F, and a G-M Laboratories portable galvanometer readily detected this difference. 2. Power The main heater power was furnished by a 7 KW Variac and the guard heater power by a 1.2KW Variac. A standard high quality calibrated voltmeter and ammeter measured the power to the main heater. It was necessary to use an auxiliary voltmeter to measure the voltage at the heater terminals on the test vessel because of the voltage drop across the carbon brushes in the slip-ring assembly. Obviously this was done only during nonrotation, During a particular run the electric current to the heater was maintained at a given value, but fluctuations in line voltage required periodic adjustment of the variac. No attempt was made to measure guard heater power.

3o RPM of Main Shaft A tachometer - generator, seen in Figure 16, generating 60 pulses per revolution was geared to the main shaft with a l:2 ratio,, Its output was fed to a Model 522B Hetlett-Packard electronic coiunterS Figure 17 is an overall view of the test apparatus and instrumentationo

III. ESTIMATION OF ERRORS A. Temperature 1. Potentiometer The manufacturer of the potentiometer (L & N, K-3) states that the maximum error at the range used is + (.15% + 0.5 v). Used in conjunction with the potentiometer were a certified standard cell and a mirror-type galvanometer with a sensitivity of 0.0003jA/mm. In order to determine the reproducibility of readings with the potentiometer, a series of 20 successive independent measurements were made with a thermocouple at the steam point in a hypsometer. The standard deviation of these readings was found to be 0.3iv, less than the maximum error given by the manufacturer. With copper-constantan thermocouples the standard deviation amounts to approximately 0.010F. In tests run at the higher values of heat flux, a tendency for the surface temperature to vary periodically with time was noted. Readings of the potentiometer were taken at the mid-point of the oscillation, and the extent of the deviation on the galvanometer scale recorded, To convert this deviation to a temperature, the deflection of the galvanometer was calibrated and found to correspond to 1.74~t volts per mm of deflection. 2. Thermocouples All thermocouples were constructed from the same spools of wire and calibrated at the steam point, and at other temperatures by comparison with Bureau of Standards calibrated mercury-in-glass thermometers. The thermocouples agreed within + liv. No difference in calibration could be detected between the water temperature thermocouples

and the heater block thermocouples made with the 1/32 inch OD. copper tubing serving as one of the conductors~ The composition of the copper thermocouple wire was not available, but the manufacturer of the tubing listed its composition as oxygen-free high conductivity copper - copper 99,96%, phosphorous - o0003% max. A thermal EMF arising from any inhomuo geniety between the wire and tube would be small tince the junction exists in an essentially isothermal zone. 3. Effect of circuit on accuracy ot measurements, The determination of errors in temperature measurements introduced by the complex circuit 'consisted.of a two-step process. A check of the system under static conditions (ie, - nonrotating) was made, whereby any error due to the presence of the iron wire, the kerosene compensating.bath, the.mercury and the switching arrangement may be detected, A check also was made under rotating conditions, which would then indicate the sole effect of the rotation, if no errors are detected in the first check, The first' check is absolute, the second relative. For the first check, one of the thermocouples in the test vessel was disconnected at a. junction box.on the rotating.assembly and another previously calibrated one substituted in its place, This thermocouple was again calibrated at the steam point, but the connection to the potentiometer was now made by way.of the entire test circuit as it is to be used. The measurement came within the + lpv deviation obtained with the original calibrations,

-36 -For the second check it was necessary that a stable temperature be available in the test vessel in order to compare rotating and nonrotating measurements. A number of trials indicated that the temperature of the water in the test vessel was not uniform enough over the period of time required in spite of the large thermal inertia.of the test vessel. The acceleration upon rotating caused convection currents which changed readings by as much as 8ktv (i.e. -0.3~F), hence overshadowing the error to be detected. The temperature of the copper heater cylinder,-however, was found to be sufficiently stable if no power were applied for at least the previous 48 hours. As was stated earlier, thermocouple T1, located nearest the heating surface in the center, was the only one of the four thermocouples in the heater cylinder to be measured with respect to the ice point. The remainder were used as differential thermocouples. Table I lists the designations of the mercury slip-ring channels and Table II lists the channel pairs associated with each thermocouple circuit. TABLE I LIST OF MERCURY THERMOCOUPLE SLIP-RING CHANNELS AND RADIUS FROM <L OF ROTATION Channel Radius from L Channel Radius from d No. of Rotation-In. No. of Rotation-In. 1 1 1/8 6 2 11/16 2 1 7/16 7 3 5 1 3/4 8 3 5/16 4 2 1/16 9 3 5/8 5 2 3/8 10 3 15/16

-37 -TABLE II LIST OF MERCURY CHANNEL PAIRS ASSOCIATED WITH EACH THERMOCOUPLE CIRCUIT Thermocouple Channel Pair Designation T1 2 &l1 AT(2-1) 3 & 4 AT(1-3) 3 & 5 zLT(2-4) 4 & 6 T5 2 & 7 T6 2 & 8 Readings of the differential thermocouples in the heater cylinder changed by no more than lliv from nonrotation to the maximum of 220 RPM, The influence of the rotation upon thermocouple Ti, through mercury channels 2 and 1, was determined by taking a rapid succession of readings at O RPM, quickly bringing the assembly to a given speed and taking another quick succession. Of readings, taking more readings at 0 RPM, and so on for the other speqdso This was repeated for thermocouples T5 and T6 by substituting thermocouple T1 in mercury channels 2 and 7 and 2 and 8 respectively. Figure 18 is a representative plot of the data taken with thermocouple T1 in mercury channels 2 and 7, The error which will be introduced in thermocouple T5 is thus indicated. Both the change measured in going up to speed and the change in coming down to 0 RPM are measured since they usually were not equal, and are plotted as corrections to be made to the readings.

I RPM 1.018 0-0 — 0 --- II0 = 0- 155- — 0 - 190 - O — 0-220- — 0 ---- 110- -0 0G 0 0 - L 0 o. w~~~~~ w... 0 1.014 W.ol4C. I o ~'00- THERMOCOUPLE TI 1i~~~~~. ~~~012 ~SUBSTITUTED FOR T5 IN 1.012 -MERCURY CHANNELS 2 & 7 0 4 8 12 16 20 24 28 32 TIME, MINUTES Figure 18. Typical Plot Tndicating Effect Qf Rotation Upon Thermocouple Readings

Figure 19 is a plot of the corrections to be applied to the thermocouples measured with respect to the ice point as a function of rotation. Many more data points were taken than are shown, but are not included for the sake of clarity, No net corrections were applied to T1 since the data deviated quite uniformly plus and minus. The cause of the EMF is not certain, but is believed due to the action of the earth's magnetic field on the uncompensated length of wire between two channels of a circuit, resulting in the presence of a unipole generator. The largest corrections must be applied to thermocouple T6, whose uncompensated length of wire has the greatest length and average velocity of (46) any of the circuits, Using tabulated values of local horizontal magnetic intensity and dip, the EMF generated at 220 RPM were calculated: Circuit T1 = 0.355v (vs + 2tv experimental) Circuit T5 = 3o17v (vs 3o.6v experimental) Circuit T6 = 4.06ktv (vs 9o8tv experimental) Also, by reversing the direction of rotation the polarity of the EMF was observed to change. Fultz and Nakagawa(39) state that they wired their rotating thermocouples noninductively to eliminate the. induced EMF's due to the horizontal component of the earth's magnetic field, In any case the net effect is known and corrections can be made, The mercury system is the largest source of error of any of the components in the temperature measuring system, It should be noted in Figure 19 that none of the data points lie more than 2~tv from the curve used in the reduction of data. With copper-constantan thermocouples and taking into consideration the other sources of error, it is believed conservative

-4o4 2 O X Q 0 0 0 0 t _-2 T, CHANNELS 1I 2 NO CORRECTIONS APPLIED U) 4 I z < O0 n~z~~~~ o -2 _ -4 0 T -2 2 7 -6 5' o - l I I 0 -4 -8 0 -10 T6, CHANNELS 288 -12 ROTATIONAL SPEED, RPM Figure 19. Corrections Applied to Thermocouples to Compensate for Rotation.

to state that the temperature at any thermocouple on the rotating assembly can be measured within + 0,l0~F 4, Determination of heater surface temperatures, The measurement of the difference in temperatures of T1 and T2 (Figure 5), AT(2-1), in the heater cylinder are made primarily to enable an extrapolation of temperature T1 to the heater surface, To do this the location of the bottom of the 1/32 inch diameter drilled holes must be known, With such deep holes small drills often have a tendency to deviate from a straight line. Two different techniques were used to locate these positions. Straight, solid, 1/32 inch diameter rods could be slid easily into the holes, indicating that the holes were straight, With a height gage on a surface plate the measurements were then extrapolated to the bottom of the holes, Figure 20 shows the locations of the holes, These are believed accurate within +.003 inches. As.a check, an ultrasonic transducer was applied to the bottom side of the heater. Reflections from interferences showed up as "pipst on a Reflectoscope, and by taking ratios of distances between "pips" on the screen the distance to the near side of a hole could be calculated, Because of other interferences it was possible to obtain clear readings for only one hole, This came within 0,006 inches of the value determined by the previous method, which is about the extent of the resolution of the Reflectoscope. To determine the uncertainty in the determination of the heating surface temperature by extrapolation from the interior, the procedure outlined by Kline and McClintock(40) is followed, The uncertainty interval

AX, =.063 HEATER SURFACE Tw - -X-. 364.056 TI T im-~~~~c TI IT Figure 20. Location of Heater Thermocouple Holes.

of a result which is obtained from a number of variables, each in themselves possessing an uncertainty, is given by WR=L(V wl) + (.R W + w(R 1 / (15) where Vn = independent variables WR = uncertainty interval W of the result R Wn = uncertainty interval of the independent variables. The uncertainty of each variable is specified by using a mean of the readings and an uncertainty interval (i.e., M + W). Referring to Figure 20 the heater surface temperature can be obtained from Equation (16). Tw T = [AT(2-l)] (16) w 1 x2 Using the term (Tw - T1) as the result in Equation (15) and dividing both sides by Equation (16) to put the uncertainty in terms of percentages gives W(T-T1) (WAx)2 + X2) > (2-1)2 ]/2 (17) Tw - T1 L 2 AT(2-1) For the dimensions of Figure 20, Ax1 =.063 +.020 inches (18) X.2 = o364 + o050 inches with a confidence limit' of 95%. The large values of W listed are necessary because it is not known which temperature across the diameter of the hole the thermocouples are actually detecting. The third term on the right hand side of Equation (17) is negligible in comparison to the first two, Substituting the above values into Equation (17): (TT) o 0,35 (19) Tw-T1

Values of W(Tw-T1) are calculated from typical test data and listed in Table III, TABLE III VALUES OF UNCERTAINTY W(TwTl) CALCULATED FROM TYPICAL TEST DATA Approx. q A AT(2-l) TW-Tl W4TwO1) Btu/Hr-Ft OF OF 10,000 1.2 0.2 0.1 25,000 3.2 0.5 0.2 50,000 6.2 1.0 0.3 75,000 10.2 1.7 o.6 100,000 13.4 2.53 o.8 The values seem large, but it must be pointed out that these result from the geometry of the system, that is, because of the unknown exact point of contact of the thermocouple junction with the heater block. For the series of tests reported here the thermocouples were not disturbed after the initial installation. Hence, although the absolute value of the heater surface temperature cannot be stated with any greater certainty than that listed in Table III, for purposes of comparison between different tests, the uncertainty will be given by the third term of Equation (17) which will give a value of W(Tw-T1) much less than + 0.10F. Since the intent is to study the effect of acceleration upon boiling heat transfer, comparisons will be made between various tests. For purposes of comparison, then, the uncertainty

of the heater surface temperature will be governed primarily by the uncertainty of the measurement T1, which was previously given as + O.10F. 5. Determination of fluid temperatures. As described earlier, the thermocouples used for measuring the water temperature are encased in stainless steel tubes which are supported from the cover of the test vessel. The cover is at ambient temperature because of the cooling water flowo The thermocouple will not give a true indication of the water temperature if conduction heat transfer up the tube wall is significant. Since the conditions present cannot be specified precisely, an error will be determined analytically by assuming the most adverse conditionspossibleo Figure 21 is a sketch showing the principle dimensions and temperatures for T6, which has the largest cross sectional area and is immersed the least in the water, The solution to the fin-type differential equation applied to this particular problem, derived in Appendix A is tw- t.2(M-K)+ K (20) memi (l+B) m(tanh me + B) m m where, twi = water temperature at the tip of the tube ta = temperature of the tube at the tip M = temperature gradient in the tube at the liquid surface K = saturation temperature gradient in the water I hC m = v~ kA = depth of immersion of the tube in the liquid hAl B = - kA

-46 -tambient = 700 F COVER. --- -I/ DIA V tw02110 F (SATURATION) t =216.5~ F (SATURATION). = THERMOCOUPLE T6 HEATER SURFACE Figure 21. Typical Conditions for Thermocouple T6 with System Subjected to Acceleration 21.15 Times Normal Gravity.

A = rosssectional area in the body of the tube A1 = cross sectional area at the closed o f the tube h = coefficient of convective heat transfer between the fluid and the fin. Making the condition even more severe, the fin is taken to be a solid rod instead of a hollow cylinder. Then A = A1, and B = h/k. Also it is assumed: h = 100 Btu/hr-ft2 k = 8 Btu/hr-ft-~F (18-8 s.s.) The largest value of K will occur at the maximum acceleration of the test chamber. At 21.15 times standard gravitational acceleration in the apparatus: K = 26.4 ~F/ft This assumes that the liquid is saturaged at all depths, likewise the most severe gradient which can exist. Assuming that the rod in the vapor space is insulated, the Temperature gradient in the rod at the water surface can be approximated by: two t amb 5640F/ft. AL 0. 25 For the 1/8 inch diameter solid rod: A = 8.5 x 10-5 ft2 C = 0.0327 ft, = 1 1/4 inches = 0.104 ft.

then m = 69.4 mD = 7.21 B = 12.5 Substituting the appropriate values into Equation (20): Error = t-t 564-2 2 + 264 (21) =69..,69.4 69.4 =.01 +.32 = o.33~F It should be noted that the second term of Equation (20) contributes the greatest part of the error. Experimental measurements of T5 and T6 in a line perpendicular to the heating surface have shown that virtually no temperature gradient exists in the part of the water probed, henceK = 0 and the error is negligible, Thermocouple T5 has a smaller diameter tube and is immersed to a greater depth than T6, hence its error is also negligible. Two additional factors aid in keeping the error small; the thermocouple supports are tubes rather than rods, and the portion of the tubes above the liquid level are exposed to saturated steam rather than being adiabatic. 6. Guard heater temperature difference The combination of the guard heater differential thermocouple and portable galvanometer were calibrated by the use of an auxiliary temporary thermocouple installed in a small hole in the heater guard (Figure 2) to measure its temperature, and thermocouple T1 to measure the temperature of the heater block. With no water in the test vessel, a small current was passed through the primary heater

-49 -and simultaneous readings taken of thermocouple T1, the auxiliary thermocouple and the portable galvanometer, It was found that one millimeter deflection of the galvanometer corresponded to a 5OF temperature differential. B. Rotation of the Main Shaft The tachometer generator geared to the main shaft generates 120 pulses for each revolution of the shaft. With the electronic counter set with a one second gate time the direct reading was twice the actual RPM, The gate time is calibrated against an internal standard crystal. Since the direct readings are accurate within + ]/2 unit, the rotational speed of the main shaft is accurate within + 1/4 RPM, During operation, the rotational speed did not vary more than 1/2 RPM over a period of several hours because of the low friction and large inertia of the system. C, Acceleration at the Heating.Surface In order to calculate the centrifugal acceleration at the heating surface the angle which the test vessel assumes must be known, It would be both difficult and inconvenient to measure this angle with any precision. Since the restraints on the pivoted test vessel were maintained at a minimum, it was deemed sufficient to calculate the angle, which requires a knowledge of the location of the center of gravity. The vessel is symmetrical about its own axis, and the axial location of the center of gravity is determined by calculating that of the counterweight. The vessel and counterweight differ in mass within

1 ounce out of approximately 30 pounds for each unit, During balancing of the unit, it was noted that moving the counterweight 0.03 inches along the threaded rod was sufficient to cause vibration at the highest rotational speed of 220 RPM. Having a simple geometrical shape, its center of gravity is readily calculated and must be.equal to that of the test vessel, within an uncertainty of + 0.03 inches. Figure 22a shows the dimensions of the test vessel necessary for calculating the acceleration at the heating surface. Referring to Figure 22b for notation, by elementary mechanics the angle 9 can be obtained from tan = W2 (b + I sin a) (22) g From trial and error solutions, 9 is listed as a function of w in Table IV for the rotational speeds commonly used in the tests. The assumption is made that the center of gravity of the assembly does not change due to the water surface taking the form of a portion of a parabola.of revolution. This is justified since the mass of the water is approximately 2 1/4 pounds as against 28 pounds for the remainder of the vessel. The centrifugal acceleration component ac at the heater surface is given by ac = w2 (b + 9H sin 9) (23) The total acceleration is the vector sum of Equation (23) and the gravitational acceleration, which is 32.17 ft/sec2 at the local latitude of 450: at = aC + g (24)

-51 -COUNTERWEIGHT - HEATING SURFACE a - PHYSICAL REPRESENTATION bC A/ REPRESEN. b - SC HEMATI C REPRESENTAT ION Figure 22. Test Vessel Dimensions for Calculation of Acceleration at the Heating Surface.

TABLE IV VALUES OF PARAMETERS IN TEST VESSEL WHICH VARY WITH ROTATIONAL SPEED Column 1. Rotational speed Column 2. Angle between heater surface and horizontal Column 3. Dimensionless total acceleration of heater surface Column 4, Angle between heater surface and acceleration vector Column 5, Total head of liquid at heater surface Column 6. Difference between angle of heater surface and liquid surface 1 2 3 4 5 6 X0$ 9 Total Acceler- Y h RPM Degrees ation aT/g Degrees Inches 9 - CP of water Degrees 0 0 1.00 90.0o 2.5 0.0 65 58.8 1.95 89.7 4.6 1.9 85 71.5 3.21 89.7 7.5 1.3 110 78.9 5.29 89.8 12.2 o.8 155 84.4 10.47 89.9 24.1 0.4 190 86.3 1573 90. o0 36.2 0. 3 220 87.2 21.15 90.0 48.4 0.2 The angle y which the total acceleration makes with the heating surface is y = $ + arctan g (25) ac Values of at/g and y also are listed in Table IV. Applying the uncertainty relation Equation (15), to the centrifugal acceleration given by Equation (23), we have Wac = 4 W.2+ C sin s n 2 H COS 3 2 1/2 as 0X b + snH + H + HCsi W (26) c W _b + 1H sin $- + +H sin $

Take as the uncertainties of the variables: W = +1/4 RPM = + 0~0262 rad/sec WQH = + 0.03 inches = + 0Q0025 fto (27) W = + 20 = 0~0349 radians Substituting into Equation (26)~ For c = 65 RPM, the uncertainty in the calculation of the acceleration at the heating surface is Wat 6 6 61/2 at = (59 3 x 10 + 6xlo + 336 x 106 + 146 x o06) 1/ 144% (28) at For w = 220 RPM, i_ =- (5.16 x 10 6 + 3076 x 106 + 1o09 x 106 032 (29) at Do Pressure on the Heating Surface With nucleate boiling taking place from a heated surface, one of the established parameters is the saturation temperature at the sure face, which can be determined if the pressure is known, For a given depth of lquid, the pressure is easily determined if the acceleration normal to the surface is linear. However, with an acceleration produced by rotation, the free surface of the liquid in Figure 22a, for example, will take on the form of a parabola of revolution, thereby complicating the calculation of the pressure at the heating surfaces Moreover, the pressure may not be uniform across the heating surface. These aspects will now be considered.

Figure 23 shows the parabolic shape of the liquid due to rotation together with that portion in the test vessel indicated by the dotted lines. Using the notation of Figures 22 and 23 the head of liquid above the center of the heater surface is h = Y2 - Yl + d cos 9 (30) Substituting the relations 2 2 2 2g o~2r2 cX xl (31) x1 = r - d sin G (32) r = b + eH sin 9 (33) into Equation (30): 2 h = [2b sin 9 + (2eH - d) sin2 e] + d cos 9 (34) h = 2 —g (3H) The assumption is made above that the depth of the liquid at the center of the test vessel does not change between the stationary and rotating conditions. Actually, the volume of liquid within the vessel is constant, and the determination of the true centerline depth would require an integration between the boundaries of the container and the liquid surface. The integral must be constant and presumably the precise location of the liquid surface could be determined as a function of rotational speed. The assumption of constant depth of liquid is justified when the difference between the heater surface and liquid surface angles is examined. Referring again to the notation of Figures 22 and 23, the angle of the liquid surface cp at the vessel centerline with respect to

-55-!1W LIQUID SURFACE V1-~~ XI _Rotation. Rotation'.

the horizontal can be determined from 2 2 tan c = g = - (r-d sin e) (35) The angle of the heater surface if given by 2 tan = (b + R sin G) (36) g and the difference, substituting Equation (33), is 2 tan G - tan cp = sin G (J + d - aH) (37) g from which ( - cp) can be determined, With the values of 9 existing, ( - cp) is a weak function of d, which justifies its use as a constant in Equation (37), Values of h as determined by Equation (34) and ($ - cp) are listed in columns 5 and 6 of Table IV for various RBM's with liquid depth d = 2,5 inches. The values of (e - cp) are largest at the lowest rotational speeds but the total liquid heads h are small, so the net effect of the error due to assuming constant liquid depth will be small. At higher rotational speeds the difference in angles (0 - p) becomes negligible, By an analysis similar to the above, it can be shown that the difference in head between the upper and lower edge of the heated surface due to the curvature of the liquid surface is given by AL = D[- cos 9 (b + iH sin 0) - sin 9] (38) where D = diameter of heater surface. At w of 220 RPM, Ah is equal to 0015 inches of water, which is negligible.

-57 -Values of the pressure and saturation temperatures at the thermocouples in the liquid are determined from expressions similar to Equation (34). An increase in pressure at the heater surface owing to acceleration of the vapor above the liquid surface is negligible. E, Heat Flux The meters used for power measurement were calibrated by comparison with precision laboratory standard instruments and found to be accurate within the resolution of the meters on all parts of the scales, Assuming that no losses exist, the heat flux is given by EKE E'I q/^A = I C x 2 (39) A r To obtain an estimate of the error in heat flux resulting solely from uncertainties in the measurements of voltage, AE, current, I, and heater radius, r, use is again made of the general uncertainty expression Equation (15). Performing the proper manipulations on Equation (39) the uncertainty of heat flux is Wq/A = r (SE) + (IWi + 4(Zr) ]l/2 (4o) q/A kE? At a heat flux of q/A; 10,000 Btu/hr-ft2, typical values are E = 31.90 + 0.05 volts I = 4.70 + 0.05 Amps. (41) r = 1.470 + 0.005 inches

Substit.uted into Equation (40) the uncertainty is w cl/A + 1-.3 (42) q/A At a heat flux of q/A \ 100,000 Btu/hr-ft2, typical values are E = 96.90 + 0.05 Volts (43) I = 14.20 + 0.05 Amps. Substituted into Equation (40) the uncertainty is: Wq/A v/A = c+ o.8% (44) The heat losses by radiation and convection from the underside of the heater are negligible because of the presence of the guard heater. A large source of error in the determination of heat flux arises from the heat loss by conduction and convection from the stainless steel heater skirt, which is in contact with the water on one face and the heater cylinder at its edge. In order to determine the order of magnitude of the loss, the skirt is treated as a straight, extended surface with a step-change in cross section. The effect of the fin being a circumferential one is negligible in this case, since the radius of curvature is large in comparison with the thickness andeffective length of the fin. Two different expressions giving the rate of heat transfer by conduction from the root of the fin for the geometry shown in Figure 4 are derived in Appendix B. One result per unit area of the main heating surface is qloss/A = 2.02 (h)l/2 9 (45)

where h = heat transfer coefficient between the fluid and the extended surface - Btu/hr-ft2-F. 00 = temperature difference between the root of the extended surface and the fluid - OF. Two tests were conducted with heat transfer to water in the convective nonboiling region, with the flow guide removed, These provided values of h for the heated surface and values of 90 for different acceleration rates, Assuming that the same h applied to the skirt, it is possible to obtain the heat loss through the skirt with Equation (45). These are tabulated in Table V along with the per cent heat loss calculated from Equation (11 The agreement between the two is noted. The percentage heat flux losses were calculated using Equation (110) for the test runs in which boiling was taking place, and are plotted in Figure 24. F, Calculation of Heat Flux Rate From Temperature Gradient in Heater Block In order to compare values of heat flux measured directly with those calculated from the temperature gradient measured in the heater block, it is essential that the uncertainty in such a calculation be determined. The heat flux is calculated from q/A k AT2) (46) Applying the uncertainty equation again: q/A [ / (4) q/A k-' +- +4

-60 -7 -2 _ 6 - ~, I I 0 5 15 20 Figure 24. /AEstima )T Loss by Conduction Through the Heater Skirt for Various3 0 5 10 15Fluxe 20 ACCELERATION - o/g Figure 24. Estimated Heat Loss by Conduction Through the Heater Skirt for Various Total Heat Fluxes.

-61 -TABLE V LOSS BY CONDUCTION THROUGH SKIRT FOR CONVECTION TESTS Test h h Heat Loss No RPM B(/A) ased on F ~ of (q/A)total llI (q/A)total | IEquatiom Equation (45) (110) C-1 0 4,700 204 23.1 13.8 12.6 110 4,700 277 16.7 11.7 11.2 155 4,700 312 15.2 11.3 11.0 190 4,700 344 13.8 10.7 10o 5 220 4,700 372 12.6 10.2 9.8 C-2 0 9,840 316 31.5 11.4 9.1 110 9,840 378 25.9 10.1 9.3 155 9,840 437 22.4 9.4 8.7

62 -The value of k given by the manufacturer for the leaded copper is the same as that for pure copper, and its reliability is uncertain. However, the uncertainty of Lx most likely will overshadow this. Taking k = 217.5 + 10 Btu/hr-ft -~F Ax =.364 + 0.05 Inches (48) AT(2-1) = 3.2 + 0.10F Wi =+ 15% (49) q/A G. Measurement of Barometric Pressure Prior to use the barometer was calibrated in the Meteorological Laboratory and corrections obtained. The barometer is accurate within + 0.005 inches of mercury. H, Measurement of Specific Resistivity of Water in the Test Vessel In order to determine the purity of the water in the test vessel, a portable conductivity cell and bridge are used to measure its specific resisticity. The instrument was checked by comparing the readings obtained with a sample of water previously measured on a precision instrument in the Physical Chemistry Laboratories of the Chemistry Department of the University. The readings agreed at one megohm within the resolution of the portable instrument, which is approximately + 5% at this range.

IV. TEST PROCEDURES A. Fluid Water was selected as the fluid to be used for this study of the effect of acceleration on boiling heat transfer. Its properties are well established and large quantities of heat transfer data are available for comparison. Double distilled water was obtained from the University's Chemistry Department and distilled again in the Heat Transfer and Thermodynamics Laboratory shortly before being used, The resistivity of the water in the test vessel was measured immediately before and after each test run. The specific resistivity of the water before each test was always 1.5 x 106 fL -cm or greater. The test vessel was filled with 1000 m2 of water, which resulted in a depth of 2.51 inches over the heater surface. For purposes of calculating the saturation temperature at the heating surface this was correctedafor the increase in specific volume with temperature. The depth.was measured again at the end of each test run. For the high heat flux test runs it did not change, but for the runs at low heat flux it was necessary to vary the cooling water flow rate to attain saturated or as near saturated conditions as possible in the water. This sometimes resulted in a net loss of vapor through the atmospheric vent. The quantity lost each time was estimated, and the total estimate prorated over the course of a run, using the total measured loss. In no case did the total change in depth exceed 0.25 inches for the test runs reported here.

-64 -After filling the test vessel and reassembling the apparatus, power was turned on and the water was boiled vigorously to degas both it and the heating surface. Prior to each test run,this was done for a minimum of 4 hours and in the majority of cases, degassing over a period of 16 hours was used. B. Heat Flux Range Covered Several test runs were made solely for obtaining convection data. These was performed with the water highly subcooled with nominal heat fluxes of 5,000 and 10,000 Btu/hr-ft2. For the test runs with boiling heat transfer the nominal heat fluxes were: 10,000 Btu/hr-ft2 25,000 Btu/hr-ft2 50,000 Btu/hr-ft2 75,000 Btu/hr-ft2 100,000 Btu/hr-ft2 For any particular test the heat flux was maintained constant while the acceleration was varied. C. Heater Surface Treatment Prior to chrominum plating, the heater surface was polished with emery cloth with successively finer grits. Finishing was done with the finest crocus cloth available. Before each lest the surface was again polished with crocus cloth and cleaned with reagent grade acetone and wiped with kleenex. After drying completely it was rinsed

twice with water of the same purity as the test water. All internal parts of the test vessel which may be in contact with either the water or vapor were subjected to the same treatment. No roughness measurements were made of the heater surface. Because.of the similar treatments above, however, it is believed that the roughness will be approximately the same for all tests. D, Range of Accelerations Covered The acceleration rates for the tests were varied from 1 to 21.15 times normal gravitation acceleration in 5 to 8 steps, In earlier tests it was noted that the wall temperatures at a/g = 1 shifted gradually over along period of time. To isolate the effect of the acceleration, the stability of a reference base was improved by taking a set of measurements at a/g = 1 before and after each higher acceleration. Thus any gradual shift in the gravitational boiling surface temperature could be compensated by considering the change only. The maximum shift.in a run which fulfilled all the requirements for acceptability was 0.80F over the average test period of 10 to 12 hours. To determine any history effect of the acceleration on the data, a sequence of accelerations similar to that given in Table VI was followed as part of the test. In most cases point numbers 2 and 6 agreed, indicating.a negligible influence of an effect of history.

TABLE VI SAMPLE SEQUENCE OF ACCELERATIONS DURING TEST RUN TO DETERMINE EFFECT OF PAST HISTORY OF ACCELERATION Point No. /.. 1 1 2 5,29 3 1 4 21.15 5 1 6 5.29,7 1 E. Location of Water Temperature Thermocouples Thermocouple T5 in Figure 2 is located on the centerline of the test vessel 0.25 inches from the heating surface for all tests. Two different thermocouples were used for T6. One was straight, as shown in Figure 2, and the other curved, permitting its location in different positions within the test vessel. Its position was varied from test to test, but for a particular run remained fixed. The purpose was to determine, if possible, the effect of the Coriolis acceleration on the flow pattern in the water, which might influence the results obtained, Figure 25 indicates the various locations used for T6, where the numbers indicate corresponding positions in the two views. F. Variation of Cooling Water Earlier tests conducted at a heat flux of 25,000 Btu/hr-ft2 with a non-restricted flow of cooling water through the condenser coils were completely non-reproducible under acceleration. It was noted that

-67 -DIRECTION OF MOTION VIEW TOWARD W2~ 1r3. 4 +HEATER SURFACE 15 __. 15 DIRECTION OF /EMOTION.__ ~~ VIEW FROM ABOVE DURING ROTATION r1 1;3 14 |x -<Ts t | | FLOW GUIDE HEATER SURFACE Figure 25. Various Locations of Thermocouple T6 for Different Test Runs.

-68 -the water had become subcooled in varying degrees, and at the highest acceleration of a/g = 21.15 it appeared likely that no boiling was taking place, even with a wall superheat of 18~F. In order to achieve reproducibility and to assure saturated or as near saturated conditions as possible it was necessary to control the cooling water flow rate for heat fluxes up to 50,000 Btu/hr-ft2. Hence it was also possible to obtain limited data on the influence of subcooling for each of the various accelerations. The subcooling occurs because of the increasingly strong convective process in the vapor space above the liquid. As will be noted in the next section, at the heat flux of 50,000 Btu/hr-ft2 variation of the cooling water had little effqct on the subcooling of the test water. The coolant flow was decreased until a net quantity of vapor passed through the atmospheric vent. At heat fluxes of 75,000 and 100,000 Btu/hr-ft2 the cooling water was not controlled. To prevent highly subcooled condensed water vapor from returning to the main body, a drip plate, shown in Figure 2, was installed over the flow guide. A number of small holes were drilled in the plate to act as a countercurrenth.eat'-exchanger, with the condensed water passing down and the vapor rising up..G Attainment of Steady State Conditions When conditions such as acceleration and cooling water flow rates were changed, sufficient time was allowed for the attainment of steady state conditions before data were taken. This ranged from 10 minutes to 1/2 hour, depending primarily upon the heat flux. At the low boiling heat flux of 10,000 Btu/hr-ft changes were sometimes so slow that pseudo-steady state data were taken.

H. Criteria for Acceptable Data For the tests reported, three conditions had to be fulfilled for acceptability of the data: 1. The specific resistivity of the test water at the conclusion of a run must be at least 800,000 -fL -cm. 2. No discolorations or spots should be present on the heating surface. 3. The value of Twall- Tsat with the system under gravitational acceleration must not change more than 0.8~F over the entire period of the test run. In almost all of the tests which were discarded, these three conditions simultaneously failed to be met, The tests performed before control of the cooling water flow was adopted also were discarded because of non-reproductibility.

V. TEST RESULTS A. Natural Convection Two tests, C-1 and C-2, were conducted with the flow guide removed at different values of heat flux in the non-boiling region. A third, Run No. C-5, was conducted with the flow guide installed to simulate the condition of heating the entire bottom surface of a container. Selected data are given in Appendix D-1. Figure 26 is a plot of the temperature difference between the heating surface and the water, Tw - T5, as a function of dimensionless acceleration a/g. The decreasein. T-T5 for a given heat flux caused by the presence of the flow guide is noted, No attempt was made to investigate the phenomena further at this time, Figure 27 shows the temperature data for Run No. C-5 as taken, as a function of time, Thermocouple T6 was located in position 2 of Figure 25, and indicated a water temperature approximately 20F higher than T5 as a result of the effect of the Coriolis acceleration on the flow pattern. The level of water temperature was not controlled, The data were correlated with the standard Nusselt-GrashofPiandtl.modi.li and are plotted in Figure 28, together with a sketch indicating the difference in configurations. Also included for reference is the correlation recommended by McAdams(l) for a horizontal heated plate facing upward. -70 -

-71 -35 30.. RUN NO. C-2, q/A a 9840 BTU/HR-FT2 NO FLOW GUIDE 25 EL 20 L \ vRUN NO. C-5, L. 20 \A 10,220 BTU/HR-FTz _<~~~ ~~ITH FLOW GUIDE 3) 15 _ - - -~RUN NO. C-l, q/A a 4700 BTU/HR-FT2 10 _ NO FLOW GUIDE 5 0 I I I I I I. 0 4 8 12 16 20 ACCELERATION, a/9 Figure 26. Plot of ATc versus Acceleration for Natural Convection Indicating the Effect of the Flow Guide.

230 210 22.. 230r 00 2- 1559RPM 190RPM 220RPM —. 20 Tw - Tsat o l n o I I0 RPM 1 RUN NO. C - 5 - q/A 10, 220 BTU/HR-FT WITH FLOW GUIDE 200u 180 0 RPTw ' 10 10 1 2 3 4 5 T IME-HOURS Figure 27. Run No. C-5. Temperature Data for Natural Convection as Taken. Convection as Taken.

-73 -0 RUN NO. C-I, q/A 4,700 BTU/ HR-FT2 16 RUN NO.C-2, q/A =9,840 BTU/ HR-FT 1000 E0 RUN NO.C-5, q/A =10,220 BTU/HR-FT 900 800 700 2 110 RPM 3 155 RPM 600 4 190 RPM 500 5 220 RPM 5 *n B0-ILING 4 400 3 0 2 z 5 200 4 13 Nu =0.14 ( Gr x Pr )13 100 RUN NO. C-I,, // C-2 RUN NO. C-5 F.LOW GUIDE \HEATER/,,I,,,,,,,,,,, 4 5 6 7 8 9 i0 2 3 4 5 6 7 8 910io Gr x Pr Figure 28. Correlation of Natural Convection with Acceleration Normal to Heating Surface.

B. Boiling For convenience and the sake of clarity, the data are.presented subdivided according to the nominal heat flux at which the tests were conducted. Except in one case, the data is representative of at least two reproducible test runs. Because of the sensitivity of the boiling process to even small degrees of subcooling at the lower values of heat flux, it was necessary to vary the water temperature in order to determine when the liquid was saturated. One of the parameters used in boiling heat transfer is the difference between the heater surface temperature and the saturation temperature of the liquid, Tw - Tsat. Because the saturation temperature varies to a large relative degree in the body of the liquid undergoing an acceleration, the saturation temperature referred to in this sense will always mean that at the heating surface. The saturation temperature of the liquid at the liquid thermocouples will also be indicated on certain plots, referring to the local saturation temperatures. 1. q/A q5 10,000 Btu/hr-ft2. The temperature data for Run No. B-15 are given in Appendix D-2. The data were carried along with the second decimal place until the last step before rounding off. Figures 29(a-d) are plots of Tw - Tsat as a function of the measured water temperature T5 for the various accelerations, The convection parameter Tw - T5 is also plotted in order to indicate when boiling has ceased. With natural convection only, Tw - T5 should be

-75 -~~~30 KNUMBERS INDICATE SEQUENCE IN WHICH 3) Tw-TSAT DATA WERE TAKEN 29 — i ---_ TwoT5 t* LOSING STEAM 28 q/A =10,870 BTU/HR-FT2 27 26 3 0 RPM 25 a/9= 4 UL 0 2- 4 5;23.7 23 TSAT@ H.S. 22 21 20 1 8 -10 21 1 717 65 RPM.s. 3 a/g-1.95 32 9; 16* TSAT ~ T TsAT @ H.S. 15 200 201 04 205 206 207 208 209 210 211 212 213 WATER TEMPERATURE, T5 OF Figure 29a. Run No. B-15. Plot of Tw-Tsat and Tw-T5 vs. Water Temperature T5 at 0 and 65 RPM.

-76 -21 2! L1t~m 13s2 I10 RPM \\ a/g =5.29 20 |3O TwT ~~~~~~20 0 ~~~~~~Tw _ Tsat 19 0 QUESTIONABLE ----— HYSTERESIS 18 -- CONVECTION \ | '' *LOSING STEAM 17 16 16 15 / / F 14/ tTsat. T5 XI abdJ~~~~~ ~Tsat. H.S. " /A =10,870 BTU/HR-FT2 155 RPM NUMBERS INDICATE SEQUENCE | IN WHICH DATA WERE TAKEN a/g =1047 3 18 I6 t ---/ — \\vl3* 14 17 -- 15 3' Tsat @ T5 13, Tsat @ H.S. 206 207 208 209 210 211 212 213 214 215 WATER TEMPERATURE T,,OF Figure 29b. Run No. B-15. Plot of Tw-Tsatandcl Tw-T5 vs. Water Temperature T5 at 110 and 155 RPM.

-77 -16 -15 14 13 `4 1 8- /10 6 14 14 9C 2 6 o. 13 / 12 12 TSAT T 5 I Ie Te13 T @RH.S. SAT9 Tw- TSAT 8 --- Tw- T5 HYSTERESIS 7 14- LOSING STEAM q/A =10,870 BTU/HR-FT 6) 190 RPM a /g = 15.73 5 NUMBERS INDICATE SEQUENCE IN WHICH DATA WERE TAKEN 205 206 207 208 209 210 211 212 213 214 215 WATER TEMPERATURE, T OF Figure 29c. Run No. B-15. Plot of Tw-Tsat and Tw-T5 vs. Water Temperature T5 at 190 RPM.

-78 -17 A,16 NUMBERS INDICATE SEQUENCE IN WHICH DATA WERE TAKEN 0 15,5 721 6 22 A 3 20 19 5 15 2, 21 I3 19 8 ITW TSATI -6 -TW- T5 7 QUESTIONABLE LOSING STEAM 16 q/A =10,8700BTU/R-FT -5 22 / 220 RPM 21a/9 21.15 206 207 208 211 212 213 214 215 216 217 218 WATER TEMPERATURE, T TOF Figure 29d. Run No. B-15. Plot of T,-T,,T and Tw-TS vs. Water Temperature T at 220 RP M Water Temperature T5 at 220 RPM.

independent of the water temperature except forchanges in liquid properties. The saturation temperatures at the heater surface and at liquid thermocouple T5 are indicated for reference. The numbers assigned to each plotted point indicate the sequence in which the data were taken. Where gaps in the sequence appear, the points were omitted from the graph to prevent overcrowding, but are listed in Appendix D-2o Points represented by ".,9 are somewhat questionable because inspection of the raw data shows that sufficient time may not have been permitted for steady-state conditions to be reached after changing the cooling water flow rate. Data points marked with an asterisk may also be somewhat questionable as these were taken while steam was issuing from the vent tube, with an attendent possible increase in pressure Referring to Figure 29a, for a/g = 1 as the subcoollng decreases the wall temperature first increases, then decreases to a minimum just prior to the point where the water has reached its saturation temperatureo For a given acceleration Tsat is constant and Tw.- Tsat is a function of the heater surface temperature Tw onlyo As the sub-.cooling decreases the contribution of the natural convection process to the total heat flux decreaseso The surface temperature then increases to provide more active sites for boiling to compensate for this decrease, A further decrease in subcooling most likely causes a thicker or more highly superheated liquid boundary layer to be formed resulting in more rapid bubble growth rates and larger bubbles which in turn result in increased agitation of the boundary layer and hence lower surf ace temperatures,

-80 For a/g = 195 the subcooling was very small, but a distinct minimum of Tw - Tsat similar to that with a/g = 1 is noted. In Figure 29b, for a/g = 5429* the wall temperature again goes through a maximum as the subcooling is decreased. only more pronounced Tw - Tsat then levels off as the saturation temperature is reached, Upon subsequent increasing of the subcooling a hysteresis effect is observed. The largest subcooling for this particular acceleration was taken at point 16, whose value of Tw - T5 is 208'~Fo This is identical to the value ofT - T5 for Run No0 C-5 in Figure 26, indicat.ing that no boiling is taking place, For a/g = 10.47, the heater surface temperature does not go through.a maximum but levels off as the sub-cooling is decreased. The value of Tw - T5 from Run Noi C-5 in Figure 26 is included as an extension of the curve of Tw - T5 to indicate the proximity of point 1 to complete non-boiling, Figure 29c for a/g =..1573 again shows a plateau of Tw - Tat with.a decrease in subcooling. A hysteresis effect also is noted, and the transition between natural convection and boiling is quite well defined. With a/g = 21,15-in Figure 29d, the transition between natural convection and boiling is nebulous, with only slight "humps" in the curves present, Points 10 and 11 are not true steady-state values as the temperatures were changing very slowly with time, The cooling water flow was completely shut off, and the high acceleration suppressed boiling until a superheat of approximately 4t.5F at thermocouple T6 was reached0 The number of active boiling sites was no

doubt very small, else the large superheat at thermocouple T5 (points 13, 14, & 15) with a net loss of steam would not be possible. The constancy of Tw - Tsat over small ranges of water temperature T5 serves as an excellent indication that well established pool boiling is taking place. The upper curve of Figure 30 is a plot of these values of Tw - Tsat as a function of acceleration, The value at a/g = 21.15 is discretionary since no well defined change in Tw - Tsat is present. The lower curves are values of the temperature differences in the heater block, with the numbers indicating the sequence of readings. For a given acceleration these values are consistant, Thermocouple T6 in the water was located in position 3 of Figure 25. Figure 31 shows.the difference in temperature indicated by the two water thermocouples T6 - T5 for the various values of T5 and acceleration. It is noted that the temperatures essentially were the same. This is also illustrated in Figure 32, where the temperatures measured between the heating surface and the liquid surface are plotted in profile for the conditions of pool boiling for the different accelerations, The numbers correspond teo the data points of Figure 29 (b-d). The local saturation temperature lines also are indicated. 2. q/a >25,000 Btu/hr-ft2. The temperature data for Run No. B-9 are given in Appendix D-3. Plots of Tw - Tsat and Tw - T5 as a function of water temperature T5 for the various accelerations are given in Figure 33 (a,b). Several items should be noted from the curves: a, The range of subcooling possible is much less than at the lower flux, which is to be expected with the cooling system used. With

-82 -19 RUN NO. B-15 ~~18|~ VA ~q/A=10,870 BTU/HR-FT l NUMBERS INDICATE 17 SEQUENCE IN WHICH 1a8 | k 3DATA WERE TAKEN i0 16 15 10 1-? 14 4 13 12 -Il8 4_ +1 0 -.H-11 02 0 0 L6I I I I I I0 0 +1 - 5il 3I 0 i 0 0 0 9 2 6 10 0 0 2 4 6 8 10 12 14 16 18 20 22 TOTAL ACCELLERATION, a /g Figure 30. Run No. B-15. Plot of Tw - Tsat and nT in Heater Block versus Acceleration with Pool Boiling.

-83 -q/A =10,870 BTU/HR-FT2 THERMOCOUPLE LOCATIONS TSAT@ T8 TSAT @ T5 T6 SA.TsT H.S 0 ' co 0220 RPM +' Ooo L O. l- i F- C0 0 0 w 155 RPM 0O, o-liO RPM +1 +I 0 p l o -O RPM 207 208 209 210 211 212 213 214 215 216 217 218 WATER TEMPERATURE, Ts 0F Figure 31. Run No. B-15. Difference in Water Temperatures for Various Accelerations and Subcooling.

-84 -234 q /A =10870 BTU/HR-F1 NOTE: NUMBERS CORRESPOND TO DATA 232 POINTS OF FIGURE 29. 2 = 220 RPM _190 RPM ao/g =21.15 228 a/g 15.73 218 216 IL ---- 214 I0o LU 0 212 TSAT SAT; 2100 w 230 9 w =155 RPM 228 3. o =1I10 RPM I O/g 5.29 o/g 10.47 226 216 214 212 -- 3 210- Io TSATTSAT 20 8 HEATER SURFACE L HEATER SURFACE WATER SURFACE WATER SURFACE Figure 32. Run No. B-15. Temperature Profile Between Heater and Water Surface for Various Accelerations.

-85 -26 25 _ 110 RPM 135 RPM a/g =5.29 a/g = 7.84 24 23 22~~~ 2 -2 _ '-4 1-2 6 6 ~~20 ~~'2 2 ~ 0 20. - 7 1-2 19 -q/A = 24,450 BTU/HR-FT2 Z 65 RPM 0W Tw - Tso t — F-ig T3 - T5 o/g =1,95 0-~- 5 QaUESTIONABLE * LOSING STEAM 24 Tsot@ T5 "[ %~H-S. Tt ~ H.S. 23 \-3 0 RPM ( 5 43 22 5 21 " 23 6 85 RPM 20 Ts- T,5 6 a/g = 3.21 209 t H.S. NUMBERS INDICATE SEQUENCE IN WHICH DATA WERE TAKE 209 210 211 212 213 214 210 211 212 213 214 215 WATER TEMPERATURE T5, OF Figure 33a. Run No. B-9. Tw-Tsat and Tw-T5 vs. Water Temperature T5 at Various Accelerations.

-86-.25,7 190 RPM 220 RPM a/g =-15.73 a/g = 21.15 23 a. \12Tsot() H.S. - 22 6 Tsot H.S. 22 Tsat ( TsS 21 -2 '3e 71,~0I ~i4: - 9 3 9 20 1 i-" 19 62 24.155 RPM q/A:24,450 BTU/ HRA ET j, 4 9 17 q/A-24,450 BTU/HR- FTZ 224 1t~T575155 RPM 21 oa/g @ 10.47 o/ -.35 2- 3 ~r QUESTIONABLE Tsat2 T \ 175 RPM 20 ~3 9 ~ Tsat ( T5 16 Ts6at( H.S. 19 9 23 765 6 18 NUMBERS INDICATE QUENCE IN WHICH DATA WERE TAKEN 17 211 212 213 214 215 212 213 214 215 216 217 WATER TEMPERATURE, T5 OF Figure 33b. Run No. B-9. TwTsat and Tw-T5 vs. Water Temperature T5 at Various Accelerations.

-87 -the heat flux input maintained constant the temperature of the water is determined by the rate of heat transfer between the surface of the water and the cooling coils above. The mechanism is a combination of liquid-water vapor interchange and natural convection..of the air in the vapor spaces The contribution of the natural convection process is essentially independent of the heat flux input. At low fluxes it is the primary mechanism while at high fluxes it is minor. As the acceleration increases its influence increases, as evidenced by the increasing range of subcooling with acceleration in Figure 33, b, In the range of subcooling covered the heater surface temperature decreases with the decrease in subcooling, indicating that not only are no new active bubble sites being formed,but the existing ones are.becoming.more effective in decreasing the thermal resistance of the boundary layer at the heating surface. This can be explained qualitatively by considering the bubbles as turbulence promotors, The turbulence induced at a single active site will be a function of both the frequency of bubble formation and the bubble volume at departure. Taking the heat transfer per bubble as proportional to the turbulence-, we have: b f Vb Of Db3 (50) Ellion(18) has shown that the average growth velocity of a bubble is

independent; of the degree of subcoolingo Neglecting any delay time between the departure of one bubble and the formation of the next, = constant (51) Since f =1/T9 Db - f = constant (52) which Jacob(2) had previously found at low values of heat fluxo If this relation holds true under accelerations greater than gravity, substitution into Equation (50) gives: )b 2 (53) Ellion (18) and Gunther(l) observed that the maximum bubble size increased with a decrease in subcoolingo According to Equation (53) then the heat transfer per bubble increases with decreased subcoolingo Since the total heat transfer rate is maintained constant, with no change in the number of nucleating sites the surface temperature will decrease, as is observed in Figure 33bo c0 At the higher accelerations it is no longer possible to attain saturated conditions near the heater surface. A temperature.profile between the heater and water surfaces for several accelerations, shown in Figure 34, aids in aceounting for this. Thermocouple T6 was located inch from the heating surface for this test run at position 1 of Figure 25, but on the basis of the previous and subsequent runs, the water temperature in a direction normal to the surface can be taken

-89 -NOTE: NUMBERS q/A 24,450 BTU/HR-Ff 236 _ | CORRESPOND TO DATA 7 7 POINTS OF FIGURE 33. 234 3I 3-2 232 23190 RPM 218 a t/g = I ' 5.29 3 12-35) 1 = 220 RPM o/g = 21.15 216 SAT 214 II~ ~ ~ ~ ~ ~ ~ ~~~,,2 ofl 2 21o LL -j 234 w=110 RPM 16 w 155 RPM Fiu232re 3 a/gR 5.29 12-3 a /g 10.47 230 21 6 214 212 210 SAT AT 208 -— ___,, HEATER SURFACE ZHEATER SURFACE WATER SURFACE WATER SURFACE Figure 34. Run No. B-9. Temperature Profile between Heater and Water Surface for Various Accelerations.

-90 -as constant up to the mid-plane between the heater and water surface and most likely closer to the water surface. When the bubble detaches from the heater surface it passes first through a locally subcooled regiony and then a more or -less superheated region. Because of the small curvature at the liquid surface, the water near the surface will not remain h:ighly superheated. The acceleration and density differences will then drive this water down toward the heating surface where it will be subcooled with respect to local conditions. A similar situation was found to exist in a study of boiling heat transfer with mercury.4 Here it was observed that the temperature throughout the bulk of the boiling mercury was essentially at the saturation temperature of the upper surface. Jacob(2 p. 623) also shows a similar disparity between local saturation and fluid temperatures with water, Figure 35 is the plot of the differences in the water temperature measured by T5 and T6, indicating that small differences do exist in a plane parallel to the heating surface. The minimum values of Tw - Tsat in Figure 33 are plotted as a function of acceleration on Figure 36, along with the temperature differences in the heater block, It may be that the small change in water temperature across the heater surface in part causes the variation of AT(1-3) with acceleration, However, the variations of AT(2-4) cannot be accounted foe, since this temperature difference is measured 7/16 inches below the heating surface, A minus value means that the centerline temperature is lower than that toward the periphery.

-91 -+l r+I TSAT 0 T5 ' fTSAT@I 0, 7~.~ q/A 24,450 BTU/HR.FT +1 00 Q — I = 190 RPM +1 0 0 035 RPM -I I =155 RPM +1 00 0 0 0 -I =6135 RPM +1 r Co O110 RPM L 210 211 212 213 214 215 216 217 THERMOCOUPLE -Figure = 85 RPM LOCATIONS +1 -I c= 65 RPM +1 z= 0 RPM Various Accelerations and Subcooling.

-92 -23 RUN NO. B-9 22 L L q/A =24,450 BTU/HR-FT2 ~ 6 NUMBERS INDICATE SEQUENCE IN WHICH DATA WERE TAKEN 21 - 20 16 19 0~~112 18 I17 16 _J 13 06 0 8 012 014 010 +1 -15-7-3 28 016 0 9-15 +1 1 37 1-3-5-9-11-15 0- - 3 06 016 82 2d ~ ~08 0I12 014 I 10 -I 4._ 16 I 17 0 82 08 012 014 010 Nm 95-9-11-13-15 0 0 P06 0 2 4 6 8 10 12 14 16 18 20 22 TOTAL ACCELERATION a /g Figure 36. Run No. B-9. Plot of Tw-Tsat and AT in Heater Block vs. Acceleration with Pool Boiling.

-93 -3. q/A 50, 000 Btu/hr-ft2o The temperature data for Run Nob B-14 are given in Appendix D-4, The range of subcooling possible with the attendant variations in Tw - Tsat were so small that the values of Tw - Tsat and water temperatures T5 and T6 are plotted directly as functions of acceleration in Figure 37. The local saturation temperatures for the water thermocouples are also included, Thermocouple T6 was located in position 3 of Figure 25 for this particular test run. The value of Tw - Tsat at a/g = 1 varied somewhat during the course of the test, and to provide a more stable reference for comparing the effects of acceleration, the average changes.ofT -Tw Tsat in coming up to and going down from a particular acceleration are shown in the lower curve of Figure 37. The variations of the differential temperatures in the heater block with acceleration are shown in Figure 38, and Figure 39 is a plot of the heater surface and water temperature profiles for several accelerations 4, q/A A 75,000 Btu/hr-ft2 The temperature data for Run No. B-22 are given in Appendix D-5* In several earlier tests at this level of rate of heat flux, attempts were made to vary the water temperature by varying the cooling water flow rate. No changes in water temperature could be detected, but undesireably large quantities of water were lost through the atmospheric vent. Subsequently no attempts were made to control the cooling water flow rate0 As an extra precaution to prevent further subcooling, however, the drip plate described earlier was installed.

-94 -28 RUN NO. B - 14 q/A 48,800 BTU/HR-FT 27 U, 25 24 NUMBERS INDICATE SEQUENCE 2217 IN WHICH DATA WERE TAKEN. 217 - 216 / - ): A H.S. U. o 215.s -- w I 6 2 -Ti1 F14 P — TS ~ ratue s) 'ra 213 THERMOCOUPLE 4 W LOCATIONS t Te 212 I10 w a 12 210 A (T - Tst) AVERAGE OF CHANGE OF T - Tt +~ IFROM 0 TO N RPM AND N TO 0 RPM. + I I- 0 94 I ~~~-~~ I 1210 0 2 4 6 8 10 12 14 16 18 20 22 TOTAL ACCELERATION, o/g Figure 37. Run No. B-14. Plot of Tw-Tsat and Water Temperatures vs. Acceleration with Pool Boiling.

-95 -+2 14 _-I, 0 014 04 <0-I - + 0 2 4 6 8 10 12 14 16 18 20 22 TOTAL ACCELERATION a/g Figure 38. Run No. 1B-14. Plot of AT in Heater Block vs. Acceleration. vs. Acceleration.

-96 -244 2 q/A 48,800 BTU/HR.FT2 24 2 240 238 _(zi= 190 RPM w = 220 RPM a/g = 15.73 a/g = 21.15 218 - 216 SAT S AT 214 - u. 212 - D210 a240 w= 110 RPM 1:55 RPM 238 - O/g 5.29 0a/g=10.47 236 216 H T TS HTE SURFC 214 R T TsAT Wae Sf 212 210 208 -_.... HEATER SURFACE RFACET WATER SURFACE — WATER SURFACEU Figure 59. Run No. B-14. Temperature Profile between Heater and Water Surface for Various Accelerations.

-97 -At this heat flux level the heater surface temperature began to oscillate with a period of between one and five seconds, The values of Tw - Tsat listed in Appendix D-5 are the mean values, and the magnitude of the variations are listed in the column following. Figure 40 is a plot of Tw - Tsat, A(Tw - Tsat) as defined previously and the differential temperatures in the heater block versus total acceleration, The entire range of accelerations were covered twice in succession, and.the data are quite reproducible in spite of a time interval of over six hours between given accelerations, Thermocouple T6 was located in position 4 of Figure 25, An earlier test, Run No, B-19, was conducted with thermocouple T6 in position 2, The heater surface temperatures were not acceptable because a number of pinpoint shadows were observed on the heater surface at the conclusion of the test and the resistivity of the water had decreased to a value of 700,000 -YL -cm, indicating a slight degree of contamination, However, it is felt that the water temperatures are not materially affected, and are plotted on Figure 41 together with those of Run No. B-22 as a function of acceleration, It should be noted that the water temperatures in the direction leading the direction of rotation are higher than those lagging, This is in the same direction that the Coriolis force acts, and hence is an indication that the flow pattern is being influenced somewhat by the Coriolis acceleration, Figure 42 is a plot of the heater surface and water temperature profiles for several representative accelerations,

-98 -RUN NO. B-22 q/A = 73,000 BTU/HR-FT 14-24 27 l1-3-7-9-11-15-23 I0,5-13-19-21 12 26 ~-~ 8 22._. 25 6-18 24 o NUMBERS INDICATE SEQUENCE IN WHICH DATA WERE TAKEN +2 24 LL. _ I -2- ~2~~~~~~ ~12 o16 r') 14 I 0 4 0 0 9 16 4 6 8 10 24 o 0 cm -8 0 0 0 24 14 -2 gure 4 R N B I I I,eate 0 01 I o N ~~~~~~~~~~~I0C~~12 10 14-24 0 2 4 6 8 10 12 14 16 18 20 22 TOTAL ACCELERATION o/g Figure 40. Run No. B-22. Plot of Tw-Tsat and AT in Heater Block vs. Acceleration with Pool Boiling.

-99 -218 RUN NO. B -22 21715 0 |q/A 73,000 BTU/HR-FT 217 > 2 16t X Tsot ( H.S 215 3 212 r 2 0 T I21 2:7 A T |q/A-73,000Tsot BTU a. 214 I - THER MOCOUPL E sa a:2 14 LOCATION W 01 O F 212 010 A12 '4 211 4 NUMBERS INDICATE SEQUENCE IN WHICH DATA WERE TAKEN 210 RUN NO. B -19 15 6I ~~q/A = 73,000 BTU/HR-FT 217 216 U. Figur 41.P Ts at (r H.S. 215 w T6 a ~5 -- Ts -I,~ /X,,__ Two (n I T < 214 a-T6 =: f w 213 THERMOCOUPLE LOCATION Aa 04 - 212 ~ 210 2 I I 210 0 2 4 6 8 10 12 14 16 18 20 22 TOTAL ACCELERATION, o/g Figure 41. Plot of Water Temperature vs. Acceleration for Two Runs Identical Except for Location of Thermocouple T6.

-100 -244 q/A:73,000 BTU/HR-FT P 242 240 W- 190 RPM W=:220 RPM o/g = 15.73 o/g 21.15 238 218 216 T T SAT S AT 214 0 212 210 - 240 =10 RPM =155 RPM o/g = 5.29 O/g = 10.47 238 236 T T6 T 6 216 214 TSAT 210 208 ~HEATER SURFACE fHEATER SURFACE WATER SURFACE -' WATER SURFACE' Figure 42. Run No. B-22. Temperature Profile Between Heater and Water Surface for Various Accelerations.

-101 -5. q/A 100,000 Btu/hr-ft2. The temperature data for Run No. B-21 are given in Appendix D-6, The oscillation of the heater surface temperatures were greater than for the previous heat flux, but decreased in magnitude with an increase in acceleration. The entire range of acceleration was again covered twice, Values of Tw - Tsat, A(Tw - Tsat) and water temperatures are shown in Figure 43 as a function of acceleration. Tw - Tsat at a/g 1 changed 0,80F during the period of the test run and except for two points, the use of A(Tw - Tsat) resulted in a better representation of the data. Figure 44 is a plot of the temperature differences within the heater block. Data points 20 and 22 of AT(1-3) are somewhat high and point 20 of AT(2-1) is somewhat low, These indicate,respectively, that the surface temperature at the center of the heater has increased and the heat transfer rate at the center had decreased, phenomena which would result from a decrease in the heat transfer coefficient at the liquid-solid interface in the center of the heater, It may be that the nucleate boiling sites were temporarily deactivated.in the vicinity of the center of the heater. Figure 45 is a plot of the surface and water temperature profiles for several accelerations, C, Overall Results 1. Boiling In spite of the precautions taken to provide consistant heater surface conditions and the purest water available for each test,

-102 -RUN NO. B- 21 q/A 99 500 BTU/HR-FT 31 20 0 LL 30 2 < >:30 i81 288 2_ + 9 19 -21- 23-26 212 28 NUMBERS INDICATE SEQUENCE I I I I HICHI DATAI WEREI TAKEN W: 215 '21 22 10 26 28 2 4 6 8 1 1- 1c~~ ~ ~~- ~ ~ ~ ~~0 -24 12 6 'F8s AT A8 216 21 5 -TOTAL ACCELERATION o/g Figure. Run No. B-21. Plot of T214at and Water Temperature vs. Acceleration with Pool - 213 - L 212 14 210 Boiling.

-103-.L o +I 20 ro G 22 7, Q2 12 14 I I., I I. I I, I 04 6 8 102 JE 2-16 1826 4 -1292 1 9 2-16 18 20 O 8 9 26 84 -I 8 24 12 8 28,I I I I i I, q/A 99,500 BTU/HR -FT2 15 NUMBERS INDICATE SEQUENCE IN WHICH DATA WERE TAKEN LL. 14~ ~12 04 9 16 I8 24 826 2 0 182 u 13 2 04 22 1,20 Q 12. II I I I I I I I I 0 2 4 6 8 10 12 14 16 18 20 22 TOTAL ACCELERATION a /g Figure 44. Run No. B-21. Plot of AT in Heater Block vs. Acceleration.

248 q/A 99,500 BTU/HR.FT 246 4 244 12 W -190 RPM = 220 RPM 242t e a/g= 15.73 a/g = 21.15 218 - 216 T-SAT TSAT 214 gW 212 S - 12 14 = 210 - a. 244 242 1 = IIORPM w= 155 RPM w ao/g - 5.29 a/g =10.47 240 - 238 214 212 210 NOTE: NUMBERS CORRESPOND TO DATA POINTS OF FIGURE 43 208 I HEATER SURFACE HEATER SUR FACE WATER SURFACE- WATER SURFACE Figure 45. Run No. B-21. Temperature Profile between Heater and Water Surface for Various Accelerations.

-105 -it was found that some of the values of Tw - Tsat with the system under gravitational acceleration were not as reproducible as was desired. In order to provide a representative composite graph.showing the influence of acceleration on Tw - Tsat for -the entire range of heat flux, the values of Tw - Tsat for a/g - 1 were plotted versus heat flux on logarithmic coordinates in Figure 46. The data for these points are given in Appendix D-7. More test runs are included than were.utilized for acceleration data because the measurements were made early in the test period while they were still considered reliable.It is noted in Appendix D-7 that even though the water had become somewhat contaminated, for the most part the measurements of Tw - Tsat at the beginning of the test period and at the end were quite reproducible, Under the influence of acceleration, however, the data became scattered. The best straight line was drawn through the points of Figure 46, and the values of Tw - Tsat from the curve were taken as the base points in plotting Tw - Tsat as a function of acceleration in Figure 4A7 for all the heat fluxes. Figure 47 thus shows the net influence that acceleration normal to a heated surface has upon pool boiling to saturated water, or rather water as nearly saturated as was possible. Test runs in addition to those presented in the previous section are included to show the degree of reproducibility. The original data points.as.well as the modified values, due to shifting the base point a/g - 1, are given in Appenfix D-8. At the low heat flux levels, an increase.in acceleration causes a decrease in Tw - Tsat. The natural convection contribution to the

66 -5 I IzJ N 4_ I0 0 / 0 1 0 10 20 30 40 TW TSAT F Figure 46. Plot of q/A vs. T-Tsat for Boiling in Standard Gravitational Field.

-107 -31 0 30 A 29 q/A 99,500 BTU/HR.FT2{ ORUN N0.B-21I -20 28 27 q/A: 73,000 BTU/HR-FT RUN NO. B-22 26 25 23 E/A 48.800 BTU/HR-FT{ 0 RUN NO. B-12 22 i A RUN NO. B-8 2 1 A q/A 24,450 BTU/HR-FT 0 s B-9 (9 FI,,,, B - I I -\ 0 " " B-Il 20 19 18 0 L 17 q/A 10,870 BTU/HR-FT 0 RUN NO. B=15 14 TOTAL ACCELERATION -o/g Figure 47. Influence of Acceleration on Tw-Tsat with Pool Boiling to Saturated Water.

-o108 -total heat transfer increases in magnitude, in effect depriving the boiling process of part of the heat flux. Fewer active nucleating sites are thus needed and the heater surface temperature decreases. It is postulated that the acceleratiorn acts to decrease the size of the bubbles at departure from the heating surface with an attendant decrease in agitation and heat transfer rate per bubble. At higher values of heat flux the increase in natural convection is not sufficient to counteract this decrease and the heater surface temperature rises to provide additional nucleating centers. At the nominal flux q/A = 100,000 Btu/hr-ft2, the surface temperature was observed to oscillate approximately ~ 0.30F with a period varying from one to five second.s, ostensibly due to the irregular boiling taking place. Upon acceleration, the magnitude of the oscillation should decrease since smaller bubbles would cause smaller irregularities. At a/g 2115 the oscillation was approximately ~ 0.1F with a period of from one to three seconds. Figure 48 is a plot of the difference between the heater surface and the water temperature, Tw - T5, corresponding to the saturated pool boiling data of Figure 47. Use of this data will be made in the next chapter in determining the contribution of natural convection to the total heat flux, The increase in Tw - T5 with acceleration at the higher values of heat flux results from the increase in subcoolingo 2, Miscellaneous Observations a- Variation of AT in heater block In analyzing the data it is cobserved that in several cases the values of the temperature differences in the heater block

-109 -38 36 q/A =99,500 BTU/HR-FT 34 _ RUN NO B-21 32 30 q/A 73,000 BTU/HR-FT| RUN NO 1- 22 28 26..| ~ —q/A =48,800 BTU/HR-FT2 RUN NO B-14 24 22 - k~/ q/A =24,450 BTU/HR-FT2?RUN NO B-9 20 8 IS \>~ <~ q/A - 10,870 BTU/HR-FT 2 16 14 12 12 I.I I I I.. 0 2 4 6 8 10 12 14 16 18 20 22 TOTAL ACCELERATION - a/g Figure 48. Plot of Tw-T5 vs. Acceleration with Pool Boiling to Saturated Water.

-110 -changed somewhat with acceleration, possibly due to non-uniform boiling conditions at the heater surface. It was at first believed to be due to the flow pattern in the water resulting from the Coriolis acceleration, but the position of the thermocouples relative to the direction of rotation was the same for all test runs, and no consistent trend is present either with.respect to heat flux nor acceleration. AT(1-3) provides a relative measure of the uniformity of the heater surface temperature, if not absolute because of the uncertainty of the exact location of the thermocouples below the surface. Where test runs were duplicated, in some cases the values of Qa(1-3) would be reproduced and in other cases they would not. The values of Tw - Tsat, however, were consistent. b, Calculation of heat transfer rate from temperature gradients in the heater The values of WT(2-1), which provide an indication of the temperature gradient at the center of the heater, were averaged for the non-rotating and the rotating conditions for each level of heat flux, The rates of heat transfer were then calculated for each of these cases and are listed in Table VII along with the percent variation from the heat flux calculated from the measurement of power input, It is recalled that the inherent uncertainty of heat flux calculated from WT(2-1) was determined to be + 15%. The larger difference at the low heat flux values may be due to the heat loss through the heater skirt, but the percentage difference decreases with rotation, whereas the heat loss through the skirt should increase with

TABLE VII COMPARISON OF HEAT FLUX CALCULATED FROM HEATER TEMPERATURE GRADIENT WITH THAT CALCULATED FROM POWER MEASUREMENT Non-Rotating Rotating Average q/A% % Run No, (Power) AT(2-1) q/A Difference AT(2-1) q/A Difference ZT(2-1) q/A Difference B-15 10,870 1,06 7,600 -30.0 1.26 9,040 -16.8 1.16 8,190 -24.6 B-7 10,580 0.96 6,880 -35.0 1.17 8,400 -20.6 1.07 7,610 -28.0 B-8 24,450 2.~45 17,600 -28.0 3.19 22,900 - 6[.3 2.82 20,500 -16.1 B-9 24,450 3.16 22,600 - 7.5 B-il 24,45p 3.14 22,500 - 8.0 B-12 48,800 5.75 41,200 -15.6 6,18 44,300 - 9.2 5.97 42,700 -12, 5 B-14 48,800 5.90 42,300 -13.3 6.13 44,000 - 9,8 6.11 43,000 -11.9 B-22 73,000 1027 10*23 10.25 73,600 + 0.8 B-20 99,500 13.65 97,900 - 1.6 13.33 95,500 - 4.0 13*49 97,000 - 2.5 B-21 99,500 135,42 96,300 - 3,2 13,23 95,000 - 4.5 13,32 95,600 - 3,9

-112 -increased acceleration, For Run No. B-15, it is noted in Figure 30 that ZkT(2-1) changes very little with acceleration0 At the low heat flux levels a small change in AT(2-1) reflects in a large change of calculated flux 0 ca Effect of Subcooling At the heat flux levels q/A = 10,000 and 25 000 Btu/hr-ft2 the sensitivity of boiling heat transfer to subcooling under the action of acceleration was demonstrated., At higher flux levels it was not possible to control the degree of subcooling with pool boiling in the system. In the series of tests reported here the water depth was maintained constant at 22 inches. By decreasing the depth of the water, presumably the subcooling would have decreased at the higher accelerations. Whether this would modify the curves of Figure 47 is still open to question. Figure 49 is a plot of the subcooling existing in the water at thermocouple T5 as a function of acceleration for the various heat flux levels. The subcooling is essentially constant for q/A > 50,000 Btu/hr-ft2 and is directly proportional to the total acceleration. For constant water depth, increasing the acceleration increases the pressure at the heating sur face, but the effect on the boiling characteristics is considered negligible in the acceleration range covered. d. Effect of Coriolis Forces The measurement of water temperatures at various locations within the system has indicated that a definite non-symnetrical temperature pattern is established in the water as a result of accelerating the system by rotation. This was attlributed to the action of the

-113 -SUBCOOLING (TsAT. T5) -T5 @ q/A 24450 BTU/HR-FT2 0 *s = 48800 o 4 + " = 73000 of h-i:,, -99500 z 1 0 0 i I I I I I 0 2 4 6 8 10 12 14 16 18 20 22 TOTAL ACCELERATION, a/g Figure 49. Subcooling at Thermocouple T5 as a Function of Heat Flux and Acceleration. C', /Vr a (CORIOLIS) = 2 w Vr Figure 50. Example of Coriolis Acceleration on a Particle Constrained to Move Radially on a Rotating Bar.

-114 -Coriolis acceleration, Figure 50 demonstrates the direction in which the acceleration acts on a particle which is constrained to move radially along a bar rotating about one end. A fluid such as water moving radially because of convection is not so constrained and the acceleration will be in the opposite direction with respect to a rotating reference line. The question of interest is; how does this effect the results obtained? Although a rigorous calculation of the component of water velocity parallel to the heating surface is not possible with the information available, an upper limit can be approximated with restrictive assumptions, This is done in Appendix C for a convection case, Run No. C-5, with q/A = 10,220 Btu/hr-ft2 and acceleration a/g = 21h15. The maximum velocity possible with the assumptions made is 0,19 ft/sec, and most likely is less than this by a factor of two or three. The.effect of this velocity is, therefore, considered negligible. The agreement of the natural convection (non-boiling) data without the flow guide in Figure 28 would appear to lend further evidence to this conclusion, With boiling, the influence of a water velocity parallel to the heating surface should be even smaller, as the intense turbulence caused by the bubbles tends to isolate the heating surface from any velocity effecto 3. Convection Correlation In Figure 28 the results of natural convection tests without the flow guide were correlated well by Nu = 0o14 (Gr Pr)l/3 (54)

However, it was noted that the flow guide caused an increase in the value of the Nusselt number for Run No. C-5. Additional convection data exist from the low heat flux boiling tests where the subcooling was increased until boiling ceased. Values of the heater surface and water temperatures were taken where Tw - T5 became independent of the water temperature, in Figure 29 for example, and the Nusselt-Grashof-Prandtl mrrroduli were calculated, These are plotted in Figure 51 with the best straight line drawn through the points. With increasing acceleration, the Nusselt number for Run No. C-5 rises somewhat above the line. This is believed due to the increase in heat loss through the heater skirt as a.result of the large subcooling existing in the water, In the remaining tests plotted, the water temperature was not far below the boiling point. The correlation Which best describes the natural convective process within the flow guide is given by the equation of the line in Figure 51: Nu =.0505 (Gr -r)096 (55)

1000 900 0 RUN NO C5 q/A 10,220 BTU/HR-FT2 800 A is B-i - T:10,870 700 0 i' s B-15 ":10,870 of + B-Il I =24,450 Sm 600 I. 0 RPM 500 2. 110 is 5 3. 155 is 4 E 4 ~ 400 4. 190 3 5. 220 300 Nu: 0.0505( Gr x Pr )O.396 200 100 - 109 1. 2 2 3 4 5 6 7 8 91010 1.5 2 3 4 Gr x Pr Figure 51. Correlation of Natural Convection with Flow Guide.

VI. ANALYSIS A. General Because of the limited quantity of data obtained and the relatively small range of conditions covered, it is believed that it would be premature to attempt a direct correlation of the data, either with existing relationships or with ones which might be derived. Further, none of the correlations in the literature which have been examined (e.g. 12,29,30,44,47,48) appear capable of following the trends shown in Figure 47, which result from acceleration of the boiling system, This may be due in part to the inadequacy and treatment of the models used to describe heat transfer by boiling, and to the neglecting of the nonboiling convection contribution. The test of the understanding of the boiling phenomenon is the ability to predict the effect of the major variables governing the process. For many applications acceleration is not a significant variable, but it must be included to complete the description of boiling since it makes the natural convective process possible. For a given geometry, say a flat horizontal heated surface, the rate of heat transfer by nonboiling natural convection is dependent upon iG and properties of the fluid. When the temperature of the surface exceeds the saturation temperature of the fluid by a certain amount the liquid generally vaporizes in discrete bubbles at the heating surface, Accompanying this is a sharp increase in heat transfer rate for a given increase in A9. Aside from the higher LZ, which will alter the properties of the fluid near the heating surface somewhat, the primary difference -117;

-1. 8 between the two processes is the formation of the bubbles which serve (2) to highly agitate the fluid, as was proposed by Jacob and others. Many of the correlations proposed for nucleate boiling initially consider the action of the bubbles in promoting the heat transfer, Because of the inability of adequately describing this action, it is always neces-, sary to revert to experiment to provide a relationship for prediction purposes. An imposed acceleration changes the buoyant forces within the fluid being heated, and any analysis of the efgect of acceleration on boiling heat transfer is really an analysis of the role of the bubbles. What will be attempted here is a presentation of the various interrelationships of bubble characteristics in boiling which have been given in the literature, together with limited extensions based on the observations made as a result of the experiments reported here. B, Bubble Relationships in Boiling Hirano and Nishikawa(49) used the simple representation of Figure 52 to point out the overall relationships involved in boiling heat transfer for the case where heat flux rate is the forcing function, as with an electrically heated surface or in a nuclear reactor. Figure 53 is a detailed extension of this, and serves to show the interdependence and complexity of the elements which are customarily considered in the boiling process. As yet little is known about the relations between the majority of these elements which must be determined before boiling can be described completely.

-119 -q/A TEMP OIL FLUID MOVE Figure 52. Representtiyn of Hirano and Nishikawa 49) for Boiling Heat Transfer. q/A A.8 BBLE \CLEVEL ANDNTION GROWmH DISTRIBUTION C NO. RATE IN ACTIVE BOUNDARY SITI Elements.

-120.'For a given system and conditions the following relationship was found to hold true over a large part of the nucleate boiling range~ q/A = C(Ag)n (56) where the interaction of the bubble elements in the lower part of Figure 53 is manifested by the values of C and n. Another form often used is q/A = h AG (57) where these interactions are now represented in the value of the convective heat transfer coefficient h. The expression developed by Fritz(14) for the maximum volume of a bubble at departure from a surface under equilibrium conditions is given in Equation (3), and good agreement was found with measurements at low values of heat flux (51) Jacob(2) experimentally determined the relation between the frequency of formation and bubble diameter at departure for saturated boiling from a horizontal surface and found f Db = Constant (58) This was also found to hold true in a statistical sense with the presence of contamination and surface active agents in water(50,51) The data were obtained from high speed photographs at low values of heat flux only, out of necessity. For methanol, Perkins and Westwater determined that f, Db and hence f Dbremained constant up to approximately 80% of the peak heat flux, after which they increased. However, a horizontal round tube was used for the heating surface, and no information was given as to where the measurements were made. Recent experimental work(3) appears to indicate that the diameter of the bubble at departure Db decreases

-121 -with an increase in heat flux rate. By using the variation in index of refraction of light in water with temperature, the thickness of a "boundacry layer" near the horizontal hieating surface with convection and pool boiling to saturated water was measured(54) at low values of heat flux. The data presented is shown in Figure 54. With values of heat flux less than 3300 Btu/hr-ft2 it was observed that hJa (59) which is a characteristic of laminar flow. In the turbulent region, where boiling is taking place the boundary layer thickness:decreases:.'more:.'sl-ow-,ly with 'increasing, heat flux. Treshchov(55) measured temperatures to within 0.0025 inches of the heating surface with high rates of heat transfer to nearly saturated water with forced convection by means of a microthermocouple, The results are shown in Figure 55, It is noted that at a distance of 0,0059 inches the temperature of the water was uninfluenced by a large increase in heat flux. By varying the surface tension with surface active agents(5 Equations (60), (61), and (62) were obtained from experimental data at low values of heat flux with saturated boiling from a flat horizontal heating surface (Maximum q/A % 25,000 Btu/hr-ft2). These relations were found to be independent, and were also found to be valid at reduced (56) pressures.

1500 FLUID -DISTILLED WATER 1000 900 " 800 I 700 U. I 600 500 400 300 q/A < 3300 BTU/HR-FTa 200 a a a a.015.02.03.04.05.06.07.08.09.1 8 -INCHES Figure 54. Thickness of Boundary Layer on a Horizontal Heating Surface as Measured by Hirano and Nishikawa(54) Using Refraction Method. 320 320, FLUID- SATURATED WATER PRESSURE - 1.2 ATM. 310 VELOCITY - 13.1 FT/SEC. 300 0El q/A O0.48 x 106BTU/HR.FT2 A a 1.07 x106 290 V U = 1.19 x106 0 a,, 1.33xl106 280 270 260 a. 250 240 230 220 lO) I 2 3 4 5 6 7 DISTANCE FROM HEATING SURFACE,INCHESxl 10 Figure 55. Effect of Heat Flux on Fluid Temperature Near Heating Surface with Forced Convection Boiling. Due to Treschov(55).

-125 -h o(N/A Dbsf)l/3 (60) A cX(N/A) -1/6 (q/A) 2/3 (61) q/A a (N/A) /2 (62) By substituting Equatiorns(60) and (61) in Equation (57), solving for q/A and equating to Equation (62), it is observed that the quantity Db3f is predicted to be independent of heat flux. A consequence of this will be enlarged upon later. Equations (61) and (62) were also valid for different degrees of roughness of the heating surface and for different.degrees of contamination of the water, the main effect of these being to change the proportionality constant. Equation (62) has been determined (53) to hold true up to heat fluxes near the burnout heat flux. As pointed out earlier, Corty and Foust(9) have shown some effects of the surface roughness on nucleate boiling, The bubble growth rate problem has been well investigated '4111) and from the agreement with experiment the solutions appear to include the important governing parameters for conditions with no subcooling. The acceleration of a boiling systnm may have a direct influence on the maximum attainable bubble size and the temperature and thickness of the superheated liquid layer near the heating surface, as indicated in Figure 53, and an indirect effect on the other elements. C. Some Observations on Boiling at Large Values of Heat Flux Equation (56) has been well established as valid over a large part of the nucleate boiling range. However, as the peak or maximum heat flux attainable with nucleate boiling is approached, distinct

departures from this relationship have been observed, some examples of which are given in Figure 56~ The point of this departure has been designated the DNB point (Departure from Nucleate Boiling) by one group(57) Defining f by (q/A)pk (6 it is noted in Figure 56 that 4 ranges from 0,52 to more than 0.78, depending on the conditions under which boiling is taking place. The inequalities are necessary because of the gaps between data points. The significance of the region between the DNB point and the peak heat flux has been given little consideration in the literature. As a bubble grows because of evaporation of liquid from the superheated boundary layer, the surrounding liquid is impelled away from the nucleating center,, In so doing, part of the liquid is moved directly away from the heating surface, and the remainder moves more or less parallel to it for a short distance, It is the rapid movement and displacement of the liquid including ejection of the bubble from the surface and liquid inflow behind, generally termed "agitation", which results in the high heat transfer rates associated with boiling. For a completely nonwetted sur(58) face Averin found that the peak heat flux was 8% of that for a wellwetted surface, and concludes that in boiling water approximately 8% of the heat transmitted by the liquid is expended on vapor formation at the surface. Rohsenow and Clark(7) estimated that the energy transport by the bubbles as "latent heat" could only account for 10% of the boiling heat transfer.

-125 -3 _ REF. CONDITIONS k(EQU. 63) O 33 VORTEX FLOW > 78.3% A 53 N - SALT SOL'N |< 56.0% X 18 V 1.I ft/sec. Subc= 50~F 2 53.0% *) ODJ em " " m Subc00OOF - 64.0% E3 63 p=- psig 2 63.0% l7 of P= 50 SOpsig > 63.0% 10 - * BURNOUT I-3 3U I CT 106 8t 10 20 30 40 50 607080 100 200 300 400 A 8 -~F Figure 56. Representative Plots of q/A vs Ao Near Peak Heat Flux.

At low values of boiling heat flux, where the nucleating,sites are relatively far apart, the heat transfer rate should be directly proportional to the number of active sites with saturated pool boiling, (2) (59) as was noted by Jacob. Zysina-Molozhen counted the number of nucleating centers with boiling water and aqueous solutions of NaCl and Glycerol for a variety of surfaces at pressures of 1-5 atmospheres, up to a maximum heat flux of 22,000 Btu/hr-ft2. The maximum number of active sites counted was 2780 per square foot~ For water, the actual number of sites at a given flux depended on the heating surface used, but a linearity between heat flux and the number of sites existed, As the number of nucleating centers increases, interactions between bubbles will occur long before they are close enough to make contact with each other, because of the displacement of the liquid. (6o) The high speed photographs of Griffith can be interpreted as indicating the type of interaction which can occur from a single nucleating site. As a result of the interactions, the bubbles no longer are able to grow to.a large size and each site becomes less effective as a fluid agitator. Hence, a proportionately Larger number of sites are required, perhaps explaining the experimental observation given by Equation (62), i.e. N/A C(q/A)2' With an increase in heat flux, it is believed that a limit is reached beyond which no further agitation of the liquid by bubbles forming at discrete sites is possible, The maximum bubble size decreases with increasing heat flux because of the increasing interference from the surrounding bubbles, and it is further believed that a minimum size exists,

dictated by the growth dynamics. The combination of this minimum size and the hydrodynamic interaction between bubbles results in the limiting effectiveness of liquid agitation owing to bubbles originating at specific sites~ This limit is tendered as being the point at which the departure from Equation (56) takes place. A minimum spacing between bubbles exists, but the surface is still far from complete vapor coverage. If an attempt is made to increase the heat flux rate beyond this point, the majority of the additional heat flux results directly in vapor generation near the heating surface, a possibility mentioned by Chang and Snyder. Bubbles no longer form at preferred active sites but occur randomly and perhaps even at very short distances away from the heating surface because of the intense turbulence. The formation of bubbles away from a surface was observed in a study of cavitation(64), which is recognized as similar in many (10) aspects to boilingo The formation of bubbles at points away from the "preferred" or active sites likewise requires proportionately higher liquid superheats0 If forming at random on the surface much higher temperatures are required to continually activate new centers, and if forming away from the surface higher superheats are required for (43) thermodynamic reasons, resulting in the bend in the plots of q/A vs AG as in Figure 56. This picture is also in keeping with the experimental observations of Gaertner and Westwater(53) where no active sites could be discerned at heat fluxes above the DNB point. At the maximum heat flux the presence of large amounts of vapor near or at the heating surface begins to impede the transfer of heat to the liquid, resulting in a decrease in heat transfer with

further increase in surface temperature. A vapor coverage of 50% at this condition has been mentioned by several workers(18,61) although it may be less, depending on geometry and test conditions, At the beginning of stable film boiling the vapor coverage is assumed to be 100%. The effect of subcooling and forced convection is to decrease the maximum size of the bubbles(184162), and should extend the heat flux at which DNB takes place because a much larger number of active sites can be sustained before the limit of bubble interaction is reached, The value of * as defined by Equation (63) may increase or decrease, depending on the relative change in (q/A)peak due to subcooling and forced convection. On the basis of the mechanism described above it might be expected that acceleration of low magnitudes will have very little influence on the heat flux at DNBo At this point the bubbles have already attained a minimum size. The action of the acceleration is to inqrease the buoyant force on the bubble, which must have some finite size before this force can act. If the growth rate to the minimum size at DNB is large, sufficient time may not be available for the buoyant forces to influence the motion of the bubble. Relations have been developed(531Y44) which predict an increase in peak heat flux with acceleration. This appears reasonable from consideration of the peak heat flux as the condition where the large quantity of vapor present near the heating surface impedes the movement of liquid toward the surface, An acceleration will aid

-129_ in the removal of vapor from the vicinity of the heating surface, permitting a higher heat flux, D, The Influence of Acceleration on Boiling Area With well developed pool boiling in a standard gravitational field, the contribution of nonboiling convection to the total heat transfer is so small that generally it is not separated from the total heat flux in analysis. With the application of an acceleration nonboiling convection can become quite appreciable, as was indicated in the test results. In:order to describe the total heat transfer, it is necessary to determine the relative contribution of the bubbles and that not resulting from the bubbles. This was accomplished in effect by Chang(30) by considering the rate of momentum exchange in the liquid near the heating surface to consist of the sum of the molecular diffusivity and the eddy diffusivity. The eddy diffusivity provided a measure of the turbulence of nonboiling convection and agitation due to the bubbles. Another method is to consider the convection and boiling contributions as separable and to weight them in some manner. This attack will be attempted here. Accordingly, the total heat transfer rate qt is q At q +B q(64) where qc is that portion due to convection and qB is that portion due to boiling, which includes the agitation induced by the bubble action, Dividing by the total area: qt/t = c/At /A (65)

If AB represents a time-averaged area projected on the heater surface in which the action of the bubbles is felt and Ac is the area in which natural convection is effective: (q/A)t = AC qB B Ac At AB At = (q/A)c(l-7) + (q/A)B 7 (66) where Y = AB/AT + Ac =At (q/A)c = qc/Ac (q/A)B = qB/A AB is not the heater surface area covered by the bubbles, but is considered differently. It might be designated as an "area of influence,. Figure 57 is an illustration of this concept. According to the previous section the agitation produced by bubbles formed at active sites on the heating surface attains a maximum value where departure from Equation (56) takes place - the DNB point. No surface area remains in which convection can be effective, and from Equation (66) 7=1 and (67) (q/A)t = (q/A)B It iS desired now to obtain values of y as a function of heat flux and acceleration~ SolvingEquation (66) for 7y (q/A)t - (q/A)c (68) 7 (68) (q/A)B - c(q/ (q/A)t is the total heat flux measured in the experiments, but means of evaluating (q/A)c and (q/A)B must be sought. (q/A)c is considered first.

-131 -01~ TIME AVERAGED PROJECTED BUBBLE AREA "AREA OF INFLUENCE" OF BUBBLES ACTIVE SITES FREE CONVECTION AREAAC SECTION OF HEATING SURFACE Figure 57. Illustration of Area of Influence of Bubble with Boiling Heat Transfer. H EATER a.WITH NO FLOW GUIDE b. WITH FLOW GUIDE Figure 58. Convective Fluid Flow Pattern in the Experimental System with and without Flow Guide.

-13.2 -Figure 28 demonstrated agreement between experimental nonboiling convection data under acceleration with the correlation Equation (54) recommended by McAdams for a horizontal keated plate facing upward. With the installation of a flow guide it was noted that for a given heat flux the temperature difference ATc decreased, or the convective heat transfer coefficient h increased. This is attributed to increased turbulence in the boundary lTyer resulting from the disruptance of the gross convection pattern present without the flow guide, as illustrated in Figure 58. It is felt that the convective pattern of Figure 58b is more analogous to the type of convection which will exist in the free convection areas outside the "area of influence" of the bubbles, in Figure 57. Hence, the expression which correlates this type of data, Equation (55), is selected for determining (q/A)c in Equation (68). The turbulence induced by an individual bubble is a function of the maximum volume of the bubble and the frequency of formation. Taking the heat transfer conjtribution of each bubble as proportional to the turbulence and approximating the bubble shape by a sphere: b = ClDb3f (69) If Ab is the area of a bubble projected on the heating surface, Abc Db2, and Ab C2f Db (70) On the basis of the discussion associated with Equation (58) f Db is taken as constant for pool boiling of a.saturated liquid, and is also

-133 -assumed to be valid under the influence of acceleration. Equation (70) then reduces to qb/ = Constant (71) for a given acceleration. It was also determined from Equations (60-62) that Db f is independent of heat flux. For a given total heat transfer area qB = Nib (72) If the "area'' of influence" of the bubbles is proportional to their area projected on the heating surface AB = C NAb (73) Dividing.Equation (72) by Equation (73) IB = b (74) AB C Ab then by Equation (71) (q/A)B = Constant (75) It is necessary to determine what this constant is as a function of acceleration. Equation (67) stated that at the DNB point that (q/A)t = (q/A)Bo No expressions are available which give this value of heat flux. As a first approximation it will be taken as equal to the peak heat flux (q/A)p. As discussed in the introduction, a number of workers have derived expressions for (q/A)p similar to Equation (11) showing an effect of acceleration, To the writer's knowledge these have not been confirmed by experiment for accelerating systems, With K2 = 0o146.Equation (11) appears to best represent experimental values of peak heat flux with pool boiling of saturated water. For the test conditions encountered here this reduces to:

-134 -(q/A) = 2.02 x 106 1/ (a/g) tu/Hr-Ft2 (76) (/p 20 v This is plotted in Figure 59 showing the influence of acceleration, Also included is the curve of Equation (10) showing the effect of a different exponent on acceleration. Values of y were calculated from Equation (68) with experimental values of (q/A)t, (q/A)c determined from Equation (55) with zTc taken from Figure 48, and (q/A)B calculated as (q/A)p from Equation (76). These are plotted in Figure 60 as a function of acceleration for the various total heat flux, and cross-plotted on logarithmic coordinates in Figure 61. The data are correlated by the expression: = (q/A) (77) where r = 14.38 + 0.3055 (a/g) s = 1.100 + 0.02273 (a/g) An interesting observation can be made from Figure 61. By extrapolating the best straight lines through the points of constant acceleration up to the value of y = 1, it is noted that at y = 1 the lines intersect at approximately common values of heat flux for accelerations a/g > 5, rather than attaining the peak heat flux values used for determining y, This states in effect that the heat flux at y = 1, the DNB point, will be independent of the acceleration, a conclusion mentioned in the previous section, The significance of this has not yet been evaluated.

-135 -1.3 T =210.5 0F FOR o/g =I SAT 1.2 1.1 EQUATION 10 LL q X:g u,, I.O::.9 -'o.8 4.7 EQUAT I ON 76 -L.666 (q/A) = F(o/g) I-.QF g r 5. E e t f A e e t o o Pe H a Flp w I.5 W 0 4.3.2 0 2 4 6 8 10 12 14 16 18 20 22 TOTAL ACCELERATION o/g Figure 59. Effect of Acceleration on Peak Heat Flux with Pool Boiling.

-136-.30.28 0 (q/A)T = 10,870 BTU/HR-FT h "I- 24, 450.26 El = " 8,800 X " - 73,000.24 0- 99,500.22.20.18.16.14 08. 06 0 2 4 6 8 10 12 14 16 18 20 22 TOTAL ACCELERATION, a/g Figure 60. Calculated Values of 7 as a Function of Total Heat Flux- and Acceleration.

-137 -1.000.900.800.700 -.600.500 - (q/A)p.400.300 -.200 - ~.100.090 0 g o/g= 1.00.080.070 =5.29.060 - = 10.47.050.X " = 15.73.040 =21.15.030 A.020 0.010.009.008.007.006.005.004.003 04.002.001 4 3 4 1567 895 2 3 4 5678906 10o 10p 2 (q/A)T- BTU/HR-FT Figure 61. Cross Plot of Figure 60.

-138 -E,o The Influence of Acceleration on the Number of Nucleating Sites Nucleation is a phenomenon which depends only upon the heating surface conditions and fluid properties, as can be concluded from the discussion in the introduction. Hence it should be affected by acceleration only insofar as the heating surface conditions and fluid properties are themselves influenced. For a given surface-liquid combination and with only small changes in pressure at the heating surface due to acceleration, the number of nucleating sites will be a function of the heating surface temperature only. If the dependence of the number of active sites on AS is known, the experimental values of L\ can be used to calculate the number of sites for the various heat fluxes and accelerations. From Equation (56), the relation between heat flux and temperature difference for a given boiling system is: q/A = l((A)n (78) Likewise, for a given system Equation (62) relates the number of active sites to heat flux: q/A = C2(N/A)1/2 (79) Equating Equations (78) and (79) and rearranging: N/A C3() 2n (80) For the system used in the experiments reported here n = 5 from Figure 460 To determine C3, the number of active sites must be known for at least one condition of q/A and A. Figure 62 is a view of the heating surface after boiling at q/A = 99n500 Btu/hr-ft2 had taken place with slightly contaminated water

-139 -Figure 62. View of Heater Surface After Test Run With Slightly Contaminated Water at Flux q/A = 99,,500 Btu/hr-ft2. (The bare portion on the left side was wiped lightly with a finger. Tw-Tsat = 29.50F, N/A = 11,100 active sites per square foot.)

-1]0.at standard gravitational acceleration for a considerable length of time. The contamination came from ceramic thermocouple tubes which were used for the measurement of water temperature during early test runs. This particular type of contamination appeared to have no effect on the temperature differences with normal gravity, and so it is possible the number of sites shown is representative of the number that would be present if no contamination existed. However, contamination from the ceramic tubes did result in nonreproducibility of data between various tests with the system under acceleration and all conditions otherwise identical. Stainless steel thermocouple protection tubes previously described were then substituted for water temperature measurements. The number of sites in Figure 62 was counted in the central portion of the heating surface and determined to be N/A = 11,100 per square foot with AS = 2905~Fo This is quite close to the value of 9,000 per square foot taken from the mean curve of (53) Gaertner and Westwater for the same heat flux, and within the experimental scatter. Substituting these values into Equation (80) we have: N/A = 0o221 x 10-10 x (AG) (81) for the system used here. Now taking values of sG from Figure 47 it is possible to determine the number of active sites as a function of heat flux and acceleration, These are plotted in Figure 63 and are similar in form to Figure 47, as is to be expected with a logarithmic plot.

-141 -q/A=99,500 BTU/HR-FT I0 q/A-73,000 BTU/HR-FT ci C) a. 3 q/A= 48,800 BTU/ HR-FT LL |C2 rq/A -24,450 BTU/HR-FTl II ( 1 q/A= 10,870 BTU/HR-FT 10 0 2 4 6 8 0 / 12 14 16 18 20 22 TOTAL ACCELERATION o/g Figure 63. Number of Nucleating Sites as a Function of Heat Flux and Acceleration.

At the lower values of heat flux acceleration increases the nonboiling convection effect, requiring fewer nucleating sites for a constant total heat fluxo At higher values of heat flux, however, the increased convection is insufficient to compensate for the decreased agitation presumably resulting from smaller bubble sizes. With an electrically heated surface then, the surface temperature must rise to provide the required additional nucleating sites. Fo Concluding Remarks A great deal of analytic and experimental work remains before boiling heat transfer will be well understood, particularly as regards the interactions between the elements which make up the boiling process as illustrated in Figure 53. Acceleration of the boiling system directs attention to the mechanism governing the departure of a bubble from the heating surface, Equation (3), due to Fritz(l), is the only relation available showing an effect of acceleration on bubble size at departure, and must be tested experimentally. The use of fluids having liquid - vapor density ratios much different from water should prove enlightening, as should the use of higher accelerations and heat flux rates up to burnout. Other variables which might be considered with an accelerating boiling system are geometry, subcooling and forced convection0

APPENDIX A DERIVATION OF EQUATION.FOR ERROR IN MEASURING WATER TEMPERATURES Referring to Figure 64, the differential equation applied to the tube as a fin is: 2 - kA (t - tw) = o (82) where: t - temperature of the tube x = distance.from surface of the liquid h = coefficient of heat transfer between the fluid and.fin c = circumference of the tube k = thermal conductivity of the tube material A - cross sectional conduction area of the tube tw = Local fluid temperature two + Kx K = Saturation temperature gradient in the liquid two= Temperature of the liquid surface Let 2 nhe tt - two m2 Substituting into Equation (82): 62_ m2 = m2Kx (83) axs whose solution is: @= c1 etx + c2 emx+ Kx (84) 145.

LIQUID SURFACE I 1 1 I X =O x=t Figure 64. Schematic of Thermocouple Tube Shown in Figure 21.

-145 -For the boundary conditions, take: (ax)x=O (a)x=0 = M = Constant (85) where: R = length of tube immersed in the liquid tw -= Temperature of the fluid at the end of the tube t = Temperature of the tube at the end B = hA1/kA A1 = Cross sectional area of the closed end of the tube Almost the same numerical result is obtained if the boundary condition -x=O = 0 is used instead of Equation (85) above. Applying the boundary conditions to Equation (84) and rearranging, the difference in temperature between the end of the tube and the local liquid temperature is: B M-K m, m K twR - tt t e1 +m B (87) mt(tanh mR + B m m Upon examining Equation (87) it is noted that as mR increases, the bracketed term on the right side tends toward zero, while the exponential increases. To determine the net effect, multiply the bracketed term by the exponential and replace the hyperbolic tangent by its exponential equivalent. Manipulation of this term results in the following: me -me B (em _ e-m) (88) mI

-146 -For m > 2.5, emg m ) e-me and Equation (88) can be written as 2 e mR(1 + B) (89) m Substituting into Equation (87) 2(m-K) K twQ - t meL Bm) m(tanhI mS +-) (90) nm m

APPENDIX B DERIVATION;OF EQUATION FOR HEAT LOSS BY CONDUCTION THROUGH HEATER SKIRT The lower surface of the skirt in Figure 4 can be considered an adiabatic surface, since the convective coefficient to air is much less than to the water on the upper surface, As an approximation, a one-dimensional temperature distribution in the skirt will be assumed. Figure 65 is a schematic representation of the skirt with the adiabatic surface on the centerline and the length wo corresponding to the 0.062 inch dimension of Figure 4. Letting: ~.= t - tf the solution to the one-dimensional flat fin problem is given by: = C1 eNX + c e-Nx (91) where.: N=\ h/ka (92) For the case of Figure 65, we have two related solutions. For 0 x Cwo 1 = C1 e1 + C2 eN (93) For Xv wo 2 = C3 eNX + C4 e (94) -147 -

t, 281 FLUID TEMP. tf Figure 65. Extended Surface Representing the Heater Skirt.

-149 -The boundary conditions are: (a) (Gl)x-O = go (b) (2)x=oo = ~ (d) (=l)x=w (2)1x=w~ By application of these boundary conditions to Equations (93) and (94) it can be shown that: 1 = (0 - 02) eNX + C2 e 1X (95) and d1 = N1 (00 - C2) e N1 C2 N1 eN1X (96) where (1+B)O% eNiwo 2 2(B cosh Nlwo + sinh NlWo) and 51 N1 61 1/2 B 2 N2 =(2 ) (98) Let L - width of the fin. The heat transfer by conduction at the base of the fin is: dO1 dO q=- kA( )x=o = - k2b1L (x i) ( o9) dx dx XOO (99)

Substituting Equations (96) and (97) with x=O, and taking one-half of the total: (1+B) eNlWo q = -k 61 LN 10o [1- B cosh N + sinh N (100) The physical dimensions from Figure 4 are: 61 =.022 inches 52.066 inches wo = 0.062 inches L = 0.785 ft. B = 0.577 The quantity within the bracket on the right side of Equation (100) was calculated for values of h of 50 and 1000 Btu/hr-ft2-F and was found to be -1.11 and -0.993 respectively. Since the values of h fall in this range it is taken as constant at -1. Thus q = k 61LN10o (101) Taking k = 8 Btu/hr-ft-~F for stainless steel and substituting the values given above, the expression for the heat loss through the skirt per unit area of the main heating surface is: loss/A = 2..02 (h) /2 (102) The above derivation assumes that h is constant. A more realistic approach is to take h as a function of Q as well as acceleration.

-151 -An approximation can be attained by using the convection correlation recommended for a horizontal surface by McAdams (1): Nu = 0.14(GrPr) (103) Solving for h: h Kg/7 (104) where Pff a 1/3 K = 0.14 kf (P2 (105) Substituting Equation (104) into the equation for the fin in Figure 65: 2 4/3 = (106) $x This is a non-linear differential equation, and a solution with the step change in cross sectional area would be quite complex. By assuming a uniform cross sectional area, Equation (106) can be integrated to: dO 6 K,7/3 + CK]1/2 (107),dX =7 k',7 Applying the boundary conditions: Ox=0 = 0 dQ (dx )x=oo =.o then C = 0. Therefore: d 6 K 7/3 1/2(108) = [= 7-/- ]

Integrating again and applying the boundary condition Gx=o = go: 1 /6 (109) _-1/6 1/2 (109) 42k5 Calculation of 0 at x = w0 of Figure 65 indicated that the increase in cross sectional area will have negligible effect on the heat loss for the conditions present. Using Equation (108), the total heat loss through the skirt per unit area of the main heating surface is: L 6 1/2 7/6 (q/A) =A ( Kk6) / (110) loss A 7 0 where L = Width of the fin A = Main heat transfer area =.0471 ft2 K = Given by Equation (105)

APPENDIX C APPROXIMATION OF VELOCITY OF FLUID AT HEATING SURFACE DUE TO CORIOLIS' FORCE Considering natural convection only the fluid is assumed to flow past the heating surface through an opening provided by the baffle shown in Figure 66, and in so doing is heated from the temperature at T5 to that at T6..For convection Run No. C-5 at a/g = 21.15: q/A = 10,220 Btu/hr-ft2 q = 471 Btu/hr. =.131 Btu/sec T6-T5 = 2 ~F The mass rate of flow through the opening is: i =q.131 =.o6O5 lbm/sec Ah 2 Any increase in the difference in temperature would decrease the mass flow rate. The velocity through the opening between the baffle and heater surface then is: m.0605 x 144 pA 60 x.75 O.19 ft/sec -155 -

-154 -DIRECTION OF ROTATION B;B$;iAFFLE = \V ' —I T HEATER SURFACE Figure 66. Model used to Calculate Maximum Possible Water Velocity Due to Coriolis Acceleration in Test Vessel.

APPENDIX D DATA -155 -

-156 -APPENDIX D-1 SELECTED NATURAL CONVECTION DATA Run No. q/A RPM a/g Tw T5 T6 T (Tw - T5) AT(2-1) LT(2-4) h C-1 4700 0 1 196.3 173.0 173.0 23.1 0.3 0.1 204 110 5.38 175.5 159.0 159.3 16.7 0.2 0 277 155 10.72 166.5 151.5 152.2 15.2 0.4 0 312 190 16.05 161.3 147.6 147.8 13.8 0.3 0 344 220 21.50 155.2 142.9 143.2 12.6 0.3 0 372 C-2 9840 0 1 227.4 195.7 196.1 31.5 0.6 0.2 316 ** 110 5.38 209.5 183.6 184.2 25.9 0.6 0.1 378 155 10.72 197.1 174.5 175.1 22.4 0.5 0 437 *C-5 10,220 0 1 228.2 201.8 201.9 26.4 1.0 0.4 388 ** 110 5.29 215.3 194.5 197.1 20.8 1.0 -0.6 493 155 10.47 202.6 185.6 188.0 17.0 1.1 -0.7 600 190 15.73 195.4 180.1 182.4 15.3 1.2 -0.8 669 220 21.15 190.8 177.0 178.6 13.8 1.2 -0.8 740 * With Flow Guide ** Boiling Taking Place Note: Accelerations for C-1 & C-2 vary from remainder of test because of modification to apparatus.

-157 -APPENDIX D-2 RUN NO. B-15 q/A = 10,870 Btu/hr-ft2 Time RPM Tw - Tsat Tw - T5 Tsat Tw T5 T6 AT(1-3) AT(2-1) AT(2-4) 9:35 0 18.9 30.0 211.2 230.0 200.0 200.0 10:05 a/g= 19.1 30.3 230.3 200.0 200.0 0.9 1.1 0:27 1 19.6 26.0 230.8 204.8 204.8:41 19.6 24.5 230.7 206.2 206.3:54 19.4 23.6 211.1 230.5 206.9 206.8 11:09 19.3 23.2 230.5 207.3 207.4:21 19.3 23.0 230.4 207.4 207.6:33 19.0 22.0 230.1 208.2 208.3:46 18.9 21.9 230.0 208.2 208.4 1.3 1.1 -0.3 12:05 18.2 19.1 229.3 210.2 210.3:14 18.5 19.1 229.6 210.5 210.6:27 18.7 19.6 229.8 210.2 210.3:42 18.5 20.1 229.6 209.5 209.7:47 18.3 19.2 229.4 210.3 210.4 1:00 18.6 18.0 229.6 211.6 211.7 *:05 19.3 18.3 230.3 212.0 212.0 *:07 19.0 18.0 230.0 212.0 2:32 / 18.5 19.0 211.0 229.5 210.4 210.4:40 110 16.9 17.8 212.2 229.1 211.3 211.3:56 a/g= 16.7 17.6 228.9 211.3 211.4 3:04 5.29 15.8 15.9 227.9 212.0 212.0:11 15.7 15.1 227.9 212.8 212.8 0.7 1.3 -0.4:17 15.7 14.7 227.8 213.2 213.2:40 15.5 14.1 227.6 213.6 213.6 *:48 15.5 14.1 227.6 213.6 213.6 4:09 15.5 15.2 212.1 227.6 212.4 212.4:21 15.9 16.3 228.1 211.8 211.8:39 16.2 16.8 228.3 211.5 211.5:57 16.9 17.7 229.0 211.3 211.4 5:13 18.1 21.2 230.3 209.1 209.2:21 18.0 21.3 230.1 208.8 208.9 5:30 17.3 21.2 229.4 208.2 208.6:38 16.5 21.1 228.6 207.4 208.0:47 15.8 20.8 227.9 207.1 207.5 xx:51 16.6 20.4 228.7 208.2 208.5:57 17.7 19.8 229.8 210.0 210.0 6:01 16.1 16.6 A m 228.2 211.6 211.6:11 0 18.6 20.0 210.9 229.5 209.6 209.8 7:11 0 18.9 19.9 210.9 229.8 209.9 210.1 1.2 1.1 -0.4 ** 7:26 220 12.9 14.4 216.3 229.2 214.8 214.8:38 a/g= 11.9 14.1 228.1 214.0 213.9:51 10.0 14.1 226.3 212.2 212.1 1.2 1.2 -0.7 **:57 21.15 10.8 13.9 227.1 213.1 213.5 8:10 13.3 14.3 229.6 215.2 215.4:19 14.2 14.6 230.5 215.8 215.9:27 14.5 14.8 230.8 216.0 216.0:32 14.9 14.7 231.2 216.5 216.4:41 14.8 15.0 231.1 216.1 216.6:45 15.7 14.5 232.0 217.6 217.6:50 16.4 14.6 232.7 218.1 218.2 *:59 14.7 14.5 231.0 216.6 216.5 * 9:01 15.3 14.4 231.5 217.1 216.9 *:05 14.9 14.8 231.2 216.4 216.5 *:08 15.1 14.7 231.4 216.7 216.8 *:11 15.0 14.7 231.3 216.5 216.6:16 14.6 14.7 230.9 216.2 216.3:30 13.3 14.3 229.6 215.2 215.2:40 12.1 14.3 228.4 214.1 214.4:50 11.4 14.1 227.6 213.5 213.4 *10. 01 9.4 14.5 225.6 211.1 210.9:18 4.6 14.2 220.8 206.6 206.8:23 0 19.3 22.7 210.9 230.2 207.5 207.6:34 19.1 21.7 210.9 230.0 208.3 208.5 1.3 1.0 0.2

-158 -APPENDIX D-2 (CONT'D) Time RPM Tw - Tsat Tw - T5 Tsat Tw T TT6 T(1-3) AT(2-1) AT(2-4) 9:22 0 19.0 21.3 211.0 230.0 208.6 208.6:36 0 18.9 21.2 211.0 229.9 208.7 208.7 1.3 1.1 -0.5:58 155 12.8 17.3 21.5 226.4 209.1 209.3 10:12 a/g= 14.0 17.8 227.6 209.8 210.1 0.9 1.3 -0.6:26 10.47 14.8 18.1 228.4 210.3 210.6:37 15.1 17.5 228.7 211.2 211.7:46 15.4 17.9 229.0 211.1 211.5:56 15.2 17.7 228.8 211.1 211.3 11:06 15.1 17.7 228.7 211.0 211.3:22 15.1 16.1 228.6 212.5 212.7: 30 15.2 15.9 228.7 212.8 213.0:46 15.4 15.7 228.9 213.2 213.4 12:00 16.4 15.2 229.9 214.8 214.6 *:06 16.3 15.2 229.9 214.6 214.7 *:09 16.2 15.1 229.8 214.7 214.7:33 15.3 17.4 / 228.8 211.4 211.6:44 0 18.6 19.6 211.0 229.6 210.0 210.0 1:11 0 18.7 20.6 211.0 229.7 209.1 209.2 1.3 1.1 -0.4:31 65 16.2 17.1 211.2 227.5 210.3 210.1:43 a/g= 16.4 17.4 227.6 210.3 209.9 1.2 1.2 -0.4 2:10 1.95 16.5 16.1 227.7 211.6 211.4:27 16.3 15.9 227.5 211.6 211.4:44 16.8 15.7 228.0 212.3 212.2:57 16.9 15.7 228.1 212.4 212.3 * 3:00 17.0 15.6 228.2 212.6 212.4 *:03 16.9 15.7 228.2 212.5 212.4:16 16.8 15.8 228.0 212.2 212.0:40 / 16.8 17.6 228.0 210.4 210.2:50 0 18.3 19.6 211.0 229.3 209.7 209.6 4:03 0 18.5 19.6 211.0 229.4 209.9 209.9 1.3 1.0 -0.4 4:23 190 13.7 15.1 214.9 228.6 213.4 213.6:539 a/g= 14.2 14.4 229.1 214.6 214.7 0.6 1.3 -0.5:48 15.73 14.1 14.3 229.0 214.7 214.8 5:06 14.2 14.3 229.1 214.8 215.1:23 14.3 13.9 229.2 215.3 215.4 *:26 14.2 13.9 229.1 215.2 215.1 *:29 14.3 14.0 \ 229.1 215.2 215.3:46 14.3 15.7 214.8 229.1 213.4 213.5 6:04 13.9 15.8 228.7 212.8 213.2:12 14.3 15.2 229.1 213.9 214.3:32 12.4 15.8 227.3 211.5 211.8:42 11.4 15.3 226.2 210.9 211.4:55 10.7 15.2 225.5 210.3 211.0 7:G8 6.7 15.3 221. 206.3 206.4:17 \ 5.2 15.3 220.0 204.7 205.0:26 0 19.4 23.8 211.0 230.4 206.6 206.9:34 0 18..8 21.4 211.0 229.8 208.4 208.4 1.3 1.0 -0.2 * Indicates losing steam ** Data questionable because of non-equilibrium, resistivity of water Note: Resistivity of water p(before test) = 1.5 X 106 JL -cm p(after test) = 0.9 x 106 _L -cm

APPENDIX D-3 RUN NO. B-9 q/A = 24,450 Btu/hr-ft2 Time RPM Tw - Tsat Tw - T5 Tsat Tw T5 T6 AT(1-3) AT(2-1) AT(2-4) 12:45 0 22.2 22.8 211.1 233.3 210.5 210.5 1:20 a/g= 22.3 22.3 233.4 211.1 211.1:31 1 22.3 22.2 233.4 211.1 211.1 0.7 2.9 0.4:39 22.2 21.9 233.3 211.4 *:42 22.3 21.7 233.4 211.7 211.5:52 22.3 22.2 233.4 211.2 211.1:57 110 20.2 21.2 212.3 232.4 211.2 211.6 2:06 a/g= 20.3 21.5 232.6 211.1 211.4 0.9 3.2 -0.2:13 5.29 20.3 21.4 232.6 211.2 211.5 *:19 20.2 20.9 232.4 211.5 212.0:24 20.1 20.5 232.3 211.8 212.2 *:29 20.2 20.2 232.4 212.2 212.6 0.7 3.4 -0.3:36 20.0 20.7 232.2 211.5 211.8:44 20.4 21.5 232.6 211.1 211.5 **:49 20.3 20.4 232.6 212.1 212.5:51 w 20.4 232.6:55 0 22.1 22.0 211.1 233.1 211.2 211.1 3:13 22.1 22.1 233.2 211.1 211.0 0.6 3.0 0.4:22 22.2 21.9 233.3 211.4 *:24 22.4 21.7 233.4 211.7:32 220 19.6 24.2 216.4 236.0 211.8 211.8:41 a/g= 19.4 23.9 235.9 211.9 212.0:47 17.5 20.7 233.9 213.2 213.1:51 21.15 17.2 20.2 233.6 213.4 213.4 0.8 3.2 -0.5 *:55 17.0 20.0 233.5 213.4 213.4 4:01 18.9 23.0 235.3 212.2:07 19.7 24.5 236.1 211.7 **:13 17.9 21.3 234.3 213.1 213.0 *:17 17.2 20.1 \ 233.6 213.5:22 0 22.2 22.0 211.1 233.3 211.3:35 0 22.2 22.2 211.1 233.3 211.1 211.0 0.7 3.1 0.3 4:41 65 21.7 21.6 211.4 233.1 211.5 212.0:47 a/g= 21.7 21.4 233.1 211.7 212.0 1.2 2.9 -0.2 *:54 1.95 22.0 21.2 233.3 212.1 212.4 5:00 22.0 21.3 233.3 212.0 x:07 21.9 21.7 233.2 211.5:21 21.9 21.8 \ 233.2 211.4 211.8:25 22.3 22.1 211.1 233.4 211.3 211.2:33 0 22.3 22.1 211.1 233.4 211.3 211.2 0.6 2.9 o.6:41 135 19.6 20.9 213.0 232.5 211.7 211.9:48 a/g= 19.6 20.9 232.6 211.7 211.9 1.1 3.3 -0.5:54 7.84 19.5 20.6 232.5 211.9 * 6:00 19.5 20.3 232.5 212.2 212.5:10 19.2 21.3 232.2 210.9 211.0:29 19.1 21.0 232.1 211.1 211.4:34 0 22.2 22.2 211.2 233.4 211.2 211.1:43 0 22.2 22.3 211.2 233.4 211.0 211.0 0.4 3.1 0.4:52 190 19.6 23.1 215.0 234.6 211.6 211.6:59 a/g= 18.1 20.8 233.1 212.3 212.4 1.0 3.2 -0.6 7:05 15.73 18.2 20.5 233.1 212.7 *:07 18.2 20.3 232.2 212.9 213.1 xx:12 18.0 20.6 232.9 212.3:18 18.7 21.7 233.7 212.0 212.0:30 20.5 24.3 235.5 211.2 211.3:40 18.6 21.4 233.6 212.2 *:46 18.2 20.3 233.2 212.9 **:49 18.0 20.5 232.9 212.5:52 18.5 21.4 ' 233.4 212.0:55 19.4 22.6 234.3 211.8:58 0 22.2 22.1 211.2 233.3 211.2 211.1 8:06 0 22.2 22.1 211.2 233.4 211.3 211.2 0.7 3.1 0.5

-160 -APPENDIX D-3 (CONT'D) Time RPM Tw - Tsat Tw - T5 Tsat Tw T5 T6 AT(1-3) AT(2-1) AT(2-4) ** 8:15 155 19.5 22.1 213.6 233.1 211.0 211.0:22 a/g= 18.8 21.0 232.5 211.5 211.7 1.1 3.3 -0.6:26 10.47 18.7 20.6 232.4 211.8 **:31 19.0 20.6 232.6 212.0:33 18.8 20.1 232.4 212.3 212.6 *:35 18.7 19.9 232.4 212.5:39 18.7 20.1 232.3 212.3:55 19.1 21.5 232.7 211.2 211.4:59 18.8 20.8 232.4 211.6 ** 9:01 19.0 20.5 232.7 212.1:03 18.8 20.1 232.5 212.3:05 18.8 20.0 232.4 212.5 *:08 18.9 19.7 232.5 212.8 213.0:12 18.9 20.2 232.5 212.3: 16 18.9 21.0 232.5 211.5:20 19.2 21.7 232.8 211.1:22 0 22.1 22.2 211.2 233.3 211.1 211.1:32 0 22.2 22.2 211.2 233.4 211.2 211.1 0.8 3.2 0.2:39 175 19.7 22.7 214.4 234.0 211.3 211.3:45 a/g= 20.0 23.2 234.3 211.2:53 18.7 21.2 233.0 211.8 211.9 1.1 3.2 -0.6:58 13.35 18.4 20.7 232.8 212.1 10:01 18.5 20.2 232.8 212.6 *:03 18.4 20.0 232.8 212.8 213.0:07 18.5 20.5 232.9 212.4:15 18.5 20.9 232.8 212.0:20 19.7 22.7 234.0 211.4:23 0 22.1 22.2 211.2 233.3 211.1 211.1:37 0 21.9 21.8 211.2 233.1 211.3 211.2 0.5 3.1 ~0.4 10:45 85 20.8 21.4 2:1.8 232.6 211.2 211.5:52 a/g= 20.7 21.1 232.5 211.4:57 321 20.6 20.9 232.4 211.5 211.8 0.5 3.5 -0.2 11:03 20.4 20.5 232.2 211.7:07 20.2 20.2 232.0 211.8 *:09 20.3 20.0 232.1 212.1 212.5 *:12 20.4 20.1 232.2 212.1:15 20.5 20.8 232.3 211.5:22 20.7 21.3 232.5 211.2 211.7:27 21.9 21.5 211.2 233.1 211.6 211.5:31 22.0 21.6 233.2 211.6:34 22.0 21.9 233.2 211.2 211.2 0.5 3.2 0.1 *:42 21.8 21.2 233.0 21.1.8: 46 21.8 22.2 233.0 210.8 210.8 * Indicates losing steam ** Data questionable because of non-equilibrium Note: Resistivity of water p(before test) = 1.4 x 106 jL -cm p(after test) = 0.8 x 106 jI -cm

APPENDIX D-4 RUN NO. B-14 q/A = 48,800 Btu/hr-ft2 Time RPM Tw - Tsat Tw - T5 Tsat Tw T5 T6 AT(1-3) AT(2-1) AT(2-4) 12:34 0 26.1 26.1 210.9 237.0 210.9 210.9 0.1 5.8 0.6:50 a/g=l 26.2 26.2 237.1 210.9 210.9 1:03 26.1 25.9 237.0 211.1 211.0:22 26.1 26.0 237.0 211.0 211.0:29 25.9 25.8 236.8 211.0 211.0 0 5.9 0.7 *:35 26.1 25.7 237.0 211.3 211.2:48 \ 26.0 25.9 236.9 211.0 210.9 2:00 110 25.0 25.8 212.1 237.1 211.3:07 a/g= 24.9 25.7 237.0 211.3 0.8 6.2 -0.1 *:17 5.29 25.1 25.5 237.2 211.7:18 25.2 25.6 237.3 211.7:25 25.1 25.9 237.2 211.3 211.5:29 25.1 25.8 237.1 211.3:35 0 26.0 25.9 210.9 236.9 211.0 210.9:46 25.9 25.9 236.8 210.9 210.9 0 5.9 0.8 3:19 4 25.9 25.9 4/ 236.8 210.9 210.8:29 220 25.8 30.0 216.3 242.0 212.0 212.1:35 a/g= 25.7 29.9 242.0 212.1:43 21.15 25.5 29.4 241.7 212.3 212.3 0.8 6.4 -0.3:54 25.4 28.9 241.6 212.7 212.6:59 25.7 28.3 242.0 213.7 213.6 * 4:00 25.7 28.3 \4 242.0 213.7 213.6:08 I 25.6 29.5 216.2 241.8 212.3 212.2:16 0 26.0 25.9 210.9 236.9 211.0 210.9:25 0 26.0 26.0 210.9 236.9 210.9 210.9:33 65 25.8 25.8 211.1 236.9 211.1 211.0:43 a/g= 25.9 25.9 237.0 211.1 210.9 1.0 5.7 0.3:50 1.95 25.9 25.7 237.0 211.3 211.2 *:52 26.0 25.7 237.2 211.5 211.2 5:00 X, 25.9 25.8 237.0 211.2 210.9:05 0 26.2 26.0 210.9 237.1 211.1 210.9:22 0 25.8 25.6 210.9 236.7 211.1 211.0 0 6.0 0.8 5:32 135 24.9 26.3 212.7 237.7 211.4 211.6:42 a/g= 25.0 26.2 237.7 211.5 211.7 0.9 6.2 0:52 7.84 24.9 25.7 237.6 211.9 212.1 *:55 25.3 26.0 238.0 212.0 212.2 6:05 25.1 26.3 237.8 211.5 211.7:10 0 25.9 25.8 O.9 236.8 211.0 210.9:39 0 26.1 26.0 210.9 237.0 211.0 210.9 0.2 5.9 0.6:51 190 25.5 28.2 214.7 240.1 211.9 211.9:57 a/g= 25.3 28.0 240.0 212.0 212.1 0.6 6.3 -0.3 7:06 15.73 25.5 27.4 240.2 212.8 212.8 *:o8 25.6 27.3 240.2 212.9 212.9:12 25.4 28.2 21h.6 240.1 211.9 212.0:17 0 26.2 26.0 210.8 237.0 211.0 210.9 10:45 0 26.2 26.1 210.8 237.1 211.0 210.9 0 6.0 0.8:57 155 25.0 26.7 213.3 238.3 211.6 211.9 11:02 a/g= 25.3 26.9 238.5 211.6:07 10.47 25.2 26.7 238.4 211.7 211.9 0.9 6.3 -0.3:14 25.0 238.3 *:16 25.4 26.3 238.7 212.4 212.6:25 24.9 26.2 238.1 211.9 212.0:28 \ 25.0 26.6 238.3 211.7 211.9:36 0 26.2 25.9 210.9 237.0 211.1 211.0:48 0 26.4 26.2 210.9 237.3 211.1 211.0 0.4 5.8 o.6:57 85 25.7 26.0 211.4 237.1 211.1 211.2 12:05 a/g= 25.6 25.8 237.0 211.2 211.3 1.2 6.0 0 *:11 3.21 25.9 25.7 237.3 211.6 211.5:16 25.7 25.9 237.1 211.2 211.2:22 0 26.1 26.0 20.9 237.0 211.0 210.9:31 0 26.1 26.0 210.9 237.0 211.0 210.9 0.1 5.9 0.6 * Indicates losing steam Note: Resistivity of water p(before test) = 1.5 x 106 fL -cm p(after test) = 0.85 x 106 j' -cm

-162 -APPENDIX D-5 RUN NO. B-22 q/A = 73,000 Btu/hr-ft2 Osc. of Time RPM Tw - Tsat Tw - Tsat Tsat Tw T5 T6 i\T(1-3) AT(2-1) AT(2-4) 2:50 0 25.7 211.0 236.7 211.2 210.9 -1.1 10.3 -0.7 3:30 0 25.6 + 0.1 211.0 236.6 211.3 211.0 -0.8 10.3 -0.8:34 0 25.6 211.0 236.7 4:20 65 25.6 211.3 236.9 211.3 211.0 -1.0 10.2 -0.9:25 65 25.6 + 0.1 211.3 236.9:39 0 25.6 211.0 236.6 211.1 210.9 -0.8 10.3 -0.8:44 0 25.6 + 0.1 211.0 236.6 5:00 85 25.4 211.6 237.0 211.3 211.1 -0.8 10.2 -1.0:04 85 25.4 211.6 237.0:16 0 25.6 211.0 236.6 211.1 210.9 -0.7 10.2 -0.7:21 0 25.7 + 0.1 211.0 236.7:47 110 25.3 + 0.1 212.2 237.5 211.4 211.2 -0.8 10.2 -1.1:52 110 25.3 212.2 237.6 6:12 0 25.6 + 0.1 211.0 236.6 211.0 210.7 -0.6 10.3 -0.8:15 0 25.7 + 0.2 211.0 236.7:29 135 25.4 213.0 238.4 211.7 211.6 -0.8 10.2 -1.1:33 135 25.3 213.0 238.3 7:15 0 25.6 + 0.1 211.0 236.7 211.2 210.9 -0.7 10.3 -0.7:19 0 25.6 + 0.1 211.0 236.7:33 155 25.5 213.7 239.1 212.0 211.7 -0.7 10.3 -1.0:39 155 25.4 ~ 0.1 213.7 239.1:54 0 25.6 211.1 236.7 211.2 211.0 -0.6 10.1 -0.8:59 0 25.7 211.1 236.7 8:16 190 26.1 + 0.0 215.1 241.2 212.4 211.8 -1.2 10.4 -0.9:21 190 26.1 215.1 241.1:37 0 25.7 211.0 236.7 211.2 211.0 -0.7 10.2 -0.6:40 0 25.7 211.0 236.7:55 220 26.9 216.5 243.4 212.6 212.1 -0.7 10.2 -1.2 9:01 220 27.0 216.5 243.5:24 0 25.6 + 0.1 211.0 236.7 211.2 211.0 -0.8 10.4 -0.7:31 0 25.6 211.0 236.6:41 65 25.5 211.3 236.9 211.3 211.1 -1.0 10.2 -1.0:46 65 25.5 + 0.1 211.3 236.8 9:57 0 25.5 211.1 236.6 211.3 211.0 10:07 110 25.3 212.3 237.6 211.6 211.4:18 0 25.7 211.1 236.8 211.4 211.1:36 155 25.4 213.7 239.1 212.0 211.8:55 0 25.7 211.1 236.8 211.3 211.0 11:11 190 25.9 215.2 241.1 212.4 211.9:27 0 25.6 ~ 0.1 211.1 236.7 211.2 211.1:48 220 27.0 216.5 243.5 212.7 212.1 -0.8 10.2 -1.1:52 220 27.0 216.5 243.5 12:06 0 25.6 + 0.2 211.1 236.6 211.2 211.0 -0.8 10.3 -o.8:11 0 25.5 + 0.2 211.1 236.6 Resistivity of water p(prior to test) = 1.35 x 106 JL -cm p(after test) = 1.0 x 106 _L -cm

-163 -APPENDIX D-6 RUN NO. B-21 q/A = 99,500 Btu/hr-ft2 Osc. of Time RPM Tw - Tsat Tw - Tsat Tsat Tw T5 T6 AT(1-3) AT(2-1) AT(2-4) 1:53 0 29.3 + 0.3 210.5 239.7 210.4 210.1 -0.5 13.7 -0.3 2:19 29.4 + 0.2 239.8 210.5 210.3:32 29.5 240.0 210.6 210.3:53 \ 29.5 + 0.3 239.9 210.4 210.3 -0.8 13.5 0 3:28 65 29.7 + 0.2 210.7 240.4 210.7 210.5 -0.2 13.2 -0.2:50 0 29.5 + 0.3 210.5 239.9 210.6 210.9 -0.9 13.7 0:54 0 29.5 + 0.2' 210.5 239.9 4:16 85 29.7 + 0.3 211.1 240.7 210.7 211.3 -0.3 13.1 -0.1:21 85 29.6 + 0.1 211.1 240.7:35 0 29.4 + 0.2 210.5 239.9 210.6 210.8 -0.8 13.6 0:40 0 29.4 + 0.1 210.5 239.9:55 110 29.7 + 0.1 211.7 241.4 210.9 211.5 0 13.1 0.2 5:00 110 29.6 + 0.1 211.7 241.3:23 0 29.4 + 0.2 210.5 239.9 210.6 210.8 -0.8 13.6 0:28 0 29.4 + 0.2 210.5 239.9:44 135 29.0 + 0.1 212.5 241.4 211.1 211.5 -0.1 13.3 -0.5:52 29.6 242.1 6:05 4 29.7 242.1 211.1 211.5:29 0 29.4 + 0.3 210.5 239.9 210.6 211.5 -0.9 13.7 0:35 0 29.3 ~ 0.2 210.5 239.8:49 155 29.9 + 0.1 213.2 243.0 211.4 212.6 0 13.2 -0.3:55 155 29.9 213.2 243.0 7:11 0 29.3 + 0.3 210.6 239.8 210.7 211.0 -0.7 13.7 -0.1:17 0 29.4 + 0.3 210.6 239.9:38 190 29.7 + 0.1 214.6 244.3 211.8 213.7 -0.1 13.5 -0.6:48 190 29.7 + 0.1 214.6 244.3 8:01 0 29.2 + 0.3 210.6 239.8 210.8 210.8 -0.9 13.7 0:07 0 29.2 + 0.3 210.6 239.8:21 220 30.5 + 0.1 216.1 246.5 212.2 213.5 0 13.8 -0.7:28 220 30.4 216.1 246.5:46 0 29.1 + 0.2 210.6 239.7 210.8 210.8 -0.9 13.6 0:55 0 29.1 210.6 239.7 9:09 65 29.0 ~+0.3 210.9 239.9 210.9 210.8 -0.8 13.3 -0.2:16 65 29.1 210.9 240.0:31 0 29.1 ~+0.3 210.7 239.8 210.8 210.9 -0.9 13.6 -0.2:37 0 29.2 210.7 239.8:52 85 29.1 ~ 0.3 211.3 240.4 210.9 211.1 -0.6 13.4 -0.2:59 85 29.1 ~ 0.2 211.3 240.4 10:16 0 29.0 210.7 239.7 210.8 210.6 -0.8 13.7 -0.1:23 0 29.0 + 0.2 210.7 239.7:35 110 30.0 211.9 241.8 211.1 211.4 0.4 12.8 -0.2:43 110 30.0 + 0.1 211.9 241.9:55 0 29.0 + 0.3 210.7 239.7 210.8 210.9 -0.9 13.8 -0.1 11:02 0 29.0 + 0.3 210.7 239.7:13 135 29.4 ~ 0.1 212.6 242.0 211.3 211.8 0.2 13.2 -0.2:21 135 29.6 212.6 242.2:27 135 29.6 0.1 212.6 242.2:40 0 29.0 + 0.3 210.7 239.7 210.9 211.1 -0.8 13.6 -0.3:46 0 29.0 + 0.2 210.7 239.7:59 155 29.4 +0.1 213.3 242.7 211.5 212.3 0.1 13.3 -0.5 12:05 155 29.4 + 0.1 213.3 242.7:20 0 29.0 t 0.3 210.7 239.7 210.8 210.8 -0.8 13.6 -0.2:26 0 29.0 + 0.3 210.7 239.7:37 190 29.4 ~0.1 214.8 244.2 212.0 213.4 -0.4 13.4 -0.5:45 190 29.4 + 0.1 214.8 244.2 1:03 0 28.7 210.7 239.4 210.8 210.7 0 13.7 - 1.7 -0.3:09 0 28.7 + 0.3 210.7 239.4:20 220 30.2 + 0.1 216.2 246.3 212.2 213.3 -0.2 13.4 -0.9:24 220 30.1 216.2 246.3:39 0 28.8 + 0.3 210.7 239.5 210.8 210.6 -0.8 13.5 -0.2:43 0 28.8 210.7 239.5 Resistivity of water p(before test) = 1.5 x 106 IL -cm p(after test) = 1.0 x 106 _L -cm

APPENDIX D-7 BOILING DATA WITH a/g = 1 FOR VARIOUS TEST RUNS Run Initial Final Initial Water Final Water Final Surface q/A No. Tw - Tsat Tw - Tsat Resistivity Resistivity Condition 10,220 C-5 18.9 19.0 10,580 B-7 17.2 17.5 1.5 x 106 _'-cm 0.9 x 106 fL-cm 10,870 B-15 18.5 18.5 1.4 x 106 0.9 x 106 24,050 B-2 22.0 22.0 1.5 x 106 0.4 x 106 Light colored spots 24,050 B-3 23.3 23.2 1.5 x 106 1.0 x 106 24,050 A-16 21.9 22.5 1.5 x 106 0.4 x 106 Light colored spots 24,050 B-6 19.7 20.1 1.5 x 106 1.0 x 106 24,450 B-8 22.0 22.2 1.5 x 106 0.8 x 106 24,450 B-9 22.2 21.9 1.4 x 106 0.8 x 106 24,450 B-11 22.4 22.5 1.4 x 106 0.9 x 106 48,800 B-12 25.3 25.2 1.4 x 106 0.8 x 106 48,800 B-14 26.0 26.1 1.5 x 106 0.9 x 106 72,700 B-17 29.5 25.7 1.5 x 106 0.6 x 106 Light colored spots 73,000 B-19 27.0 26.5 1.4 x 106 0.7 x 106 Pinpoint shadows 73,000 B-22 25.6 25.5 1.4 x 106 1.0 x 106 99,500 B-20 29.4 28.6 1.5 x 106 1.1 x 106 99,500 B-21 29.5 28.8 1.5 x 106 1.1 x 106

-165 -APPENDIX D-8 SUMMARY OF DATA USED IN FIGURE 47 Run Original Modified No. q/A RPM Tsat Tw - Tsat Tw - T5 Tw - Tsat Tw - T5 B-15 10,870 0 211.0 18.5 18.5 18.8 18.8 65 211.2 16.3 16.5 16.6 16.8 110 212.1 15.8 15.9 16.1 15.9 155 213.5 15.1 16.1 15.4 16.2 190 214.8 14.2 14.9 14.5 15.2 220 216.3 13.3 14.4 13.6 15.2 B-7 10,580 0 210.2 17.3 17.3 18.8 110 211.3 13.7 14.0 15.2 125 211.7 13.8 14.0 15.53 155 212.7 13.5 14.0 15.0 190 214.1 13.4 14.0 14.9 B-9 24,450 0 211.1 22.2 22.2 22.1 22.1 65 211.4 21.9 21.9 21.8 21.4 85 211.8 20.4 20.5 20.3 20.8 110 212.3 20.1 20.8 20.0 20.8 135 213.0 19.5 20.8 19.4 20.6 155 213.6 18.7 20.5 18.6 20.4 175 214.0 18.4 20.7 18.3 20.0 190 214.9 18.1 20.6 18.0 20.4 220 216.4 17.1 20.1 17.0 20.1 B-8 24,450 0 210.9 22.2 22.3 22.1 65 211.1 21.2 21.5 21.1 85 211.4 20.8 21.5 20.7 110 212.2 20.0 21.3 19.9 135 212.7 18.7 20.7 18.6 155 213.2 18.3 20.0 18.2 190 214.8 17.6 20.0 17.5 220 215.9 17.2 20.3 17.1 B-ll 24,450 0 210.7 22.4 22.4 22.1 65 210.9 22.7 22.7 22.4 155 213.2 19.0 19.9 18.7 220 215.8 17.5 20.1 17.2 B-12 48,800 0 209.5 25.3 25.3 25.4 65 209.9 24.8 24.8 24.6 110 210.7 24.5 25.1 24.7 135 211.5 24.6 25.9 24.5 190 213.5 24.5 27.1 24.7 220 214.8 25.0 28.8 25.1 B-14 48,800 0 210.9 26.1 26.1 25.4 25.4 65 211.1 25.9 25.9 25.3 25.1 85 211.4 25.7 25.9 24.9 25.0 110 212.1 25.0 25.8 24.4 25.3 135 212.7 25.0 26.3 24.4 25.5 155 213.3 25.1 26.8 24.3 26.0 190 214.7 25.4 28.1 24.7 27.4 220 216.3 25.5 29.3 25.0 29.0 B-22 73,000 0 211.0 25.6 25.3 27.7 27.4 65 211.3 25.6 25.6 27.7 27.6 85 211.6 25.4 25.7 27.5 27.8 110 212.2 25.3 26.1 27.4 28.2 135 213.0 25.4 26.7 27.5 28.7 155 213.7 25.4 27.1 27.5 29.2 190 215.1 26.0 28.7 28.0 30.8 220 216.5 27.0 30.9 29.1 33.0 B-21 99,500 0 210.5 29.5 29.5 29.5 29.5 65 210.7 29.7 29.7 29.7 29.5 85 211.1 29.6 30.0 29.6 30.0 110 211.7 29.7 30.5 29.8 30.4 135 212.5 29.6 31.0 29.8 31.2 155 213.2 29.9 31.7 30.1 31.7 190 214.6 29.7 32.5 30.0 32.9 220 216.1 30.5 34.4 30.8 34.7

-166 -APPENDIX D-8 (CONT'D) Run Original Modified No. q/A RPM Tsat Tw - Tsat Tw - T5 Tw - Tsat Tw - T! B-21 99,500 0 210.6 29.1 28.9 29.5 65 210.9 29.1 29.1 29.6 85 211.3 29.1 29.5 29.6 110 211.9 30.0 30.8 30.5 135 212.6 29.6 30.9 30.2 155 213.3 29.4 31.2 29.9 190 214.8 29.4 32.2 30.1 220 216.2 30.1 34.1 30.9 B-20 99,500 0 210.8 29.4 29.4 29.5 65 211.2 29.0 29.0 29.4 85 211.5 29.2 29.5 29.9 110 212.0 29.4 30.2 29.6 135 212.8 29.1 30.4 29.9 155 213.4 29.1 30.8 29.7 190 214.8 29.7 32.3 30.2 220 216.2 30.0 33.8 30.2 110 212.1 29.4 30.3 29.6 220 216.2 29.8 33.5 30.7

BIBLIOGRAPHY 1, MeAdams, W. H. Heat Transmission, 3rd Edition, New York: McGraw-Hill Book Co. Inc., 1954. 2. Jacob, M, Heat Transfer, I. New York: John Wiley & Sons, Inc.,, 1949* 3. Bosworth, R. C. L. Heat Transfer Phenomena. New York: John Wiley & Sons Inc., 1952. 4. Griffith, P. "Bubble Growth Rates in Boiling." Trans. ASME, 80, No..3, (April, 1958), 721. 5. Zuber, N. "On the Stability of Boiling Heat Transfer," Ibid, 711. 6* Westwater, J, W. and Santangelo, J. G, "Photographic Study of Boiling," Ind, & Eng. Chem., 47, (August, 1955), 1605. 7. Rohsenow, W. M, and Clark, J. A. "A Study of the Mechanism of Boiling Heat Transfer," Trans. ASME, 73, No. 5, (July, 1951), 609. 8. Clark, J. A. and Rohsenow, W. M. "Local Boiling Heat Transfer to Water at Low Reynold's Numbers and High Pressures," Trans. ASME, 76, No. 4, (1954), 553. 9. Corty, C. and Foust, A. S. "Surface Variables in Nucleate Boiling." Chem. Eng* Prog. Symp., Series No, 17, 51, 1955. 10. Clark, J. A. "Thermodynamics of Bubbles." M.I.T., Cambridge, Mass..OTIC,, Technical Report-NHo. 7, OONER Contract N5 -or i-07827 (NR-035-267), January 1l, 1956. 11. Forster, H. K. and Zuber, N. "Growth of a Vapor Bubble in a Superheated Liquid." J. App. Phys., 25, No. 4, (1954), 474* 12, Forster, H. K. and Zuber, N. "Dynamics of Vapor Bubbles and Boiling Heat Transfer." AIChE Journal, 1, No. 4, (December, 1955), 531. 135 Plesset, M. S. and Zwick, S. A. "The Growth of Vapor Bubbles in Superheated Liquids," J. App, Phys., 25, No. 4, (1954), 493. 14. Fritz, W. "Calculation of the Maximum Volumes of Vapor Bubbles." Physik, Zeitschr., 36, (1935), 379. 15. Kreith, F. and Summerfield, M. "Heat Transfer to Water at High Flux Densities With and Without Surface Boiling." Trans. ASME, 71, (1949), 805. -167 -

16. McAdams, W. H. et al, "Heat Transfer at High Rates to Water with Surface Boilings." Ind. & Eng. Chem., 41, No. 9, (September, 1949), 1945. 17. Pramuk, F. S. and Westwater, J. W. "Effect of Agitation on the Critical Temperature Difference for a Boiling Liquid."v Chemo Eng. Prog. Symposium, 52, No. 18, 1956. 18. Ellion, M. "A Study of the Mechanism of Boiling Heat Transfer.g J.P.L., Memo 20-88, Pasedena, Calif., March, 1954. 19. Insinger, T. H, Transmission of Heat to Boiling Liquids.~ PhD Dissertation, Univ. of Penn., Philadelphia, Penn., 1940. 20. Morgan, A. I., Bromley, L. A., and Wilke, C. R. "Effect of Surface Tension on Heat Transfer in Boiling." Ind. & Eng. Chem., 41, Noo 12, (December, 1949), 2676. 21. Bankoff, S. G. "Ebullition from Solid Surfaces in the Absence of a Pre-existing Gaseous Phase." Trans. ASME, 79, (May, 1957), 735. 22. Wakeshima, H. and Takada, K. "Limit of Superheat." J. Appl. Phys. 29, (July, 1938), 1126. 23. Mead, B. R., Romie, F. E.,and Guibert, A, G. "'Liquid Superheat and Boiling Heat Transfer." Heat Transfer & Fluid Mechanics Institute, Preprints of papers, 1951, Stahford University Press, Stanford, Calif. 24, Sabersky, R. H. and Gates, C. W. "Effect of Pressure on Start of Nucleation in Boiling Heat Transfer," Jet Propulsion, 25, No. 2, (February, 1955), 67. 25, Bashforth, F., and Adams, Jo Capillary Action, Cambridge, 1883. 26. Wark, J. W. "The Physical Chemistry of Flotation." Journal of Physical Chem., 37, (1933), 623. 27. Pike, F. R., Miller, P. D., and Beatty, K. 0. 'Effect of Gas EvbOlution on Surface Boiling at Wire Coils." Chem. Eng, Prog. Sympo Series, 51, No. 17, 1955. 28, Larson, R, F. "Factors that Influence Heat Transfer in Boiling." Heat Transfer and Fluid Mechanics Institute, Stanford Univbrsity Press, Stanf ford, Califf, 19535 29. Rohsenow, W, M. "A Method of Correlating Heat Transfer Data for Surface Boiling of Liquids." Trans. ASME, 74, (1952), 969.

-169.30. Chang, Y. P. "Theoretical Analysis of Heat Transfer in Natural Convection and in Boiling." Trans. ASME, 79, (October, 1957), 1501. 31. Borishanskii, Vo M. "An Equation Generalizing Experimental Data on the Cessation of Bubble Boiling in a Large Volume of Liquid." Zhurn, Tekh. Fiz,, 26, (1956), 452. Translated in Soviet PhysicsTechnical Physics, 1, No. 2, 438, Am. Inst. of Phys., N.Y. 32. -Siegel. R. and Usiskin, C. "A Photographic Study of Boiling in the Absence of Gravity." Trans. ASME, 81, Series C, No. 13, (August, 1959), 230. 335 Gambill, W. R. and Green, N. D. "A Study of Burnout Heat Fluxes Associated with Forced-Convection, Subcooled, and Bulk Nucleate Boiling of Water in Source - Vortex Flow." Chem. Eng. Prog., 54, No. 10, (October, 1958), 68. 34. Schmidt, E. H. W. "Heat Transmission by Natural Convection at High Centrifugal Acceleration in Water-Cooled Gas-Turbine Blades." The Institution of Mech, Eng. ASME, Proceedings of the General Discussion on Heat Transfer, 11-13 September, 1951. 35. Hickman, K. C. D. "Centrifugal Boiler Compression Still." Ind. & Eng. Chem,, 49, No. 5, (May, 1957), 786. 36. Gudheim, A. R. and Donovan, J.."Heat Transfer in Thin Film Centrifugal Processing Units." Chem. Eng. Prog., 53, No. 10, (October, 1957), 476. 37. Sparrow, E. M. and Gregg, J. L. "A Theory of Rotating Condensation." Trans. ASME, Journ, of Heat Transfer, 81, Series C, No. 2, (May, 1959), 1135 38. Clark, H. B., Strenge, P. S., and Westwater, J. W. "Active Sites for Nucleate Boiling." Preprint # 13 presented at AIChE-ASME 2nd National Conference on Heat Transfer, Chicago, Illinois,. August 18-21, 1958. 39. Fultz, D. and Nakagawa, Y. "Experiments on Over-Stable Thermal Convection in Mercury." Proc. Roy. Soc., (London) A 231, (1955), 211. 40. Kline, S. J. and McClintock, F. A. "Description and Analysis of Uncertainties in Single Sample Experiments," Aerodynamics Measurements, M.I.T., Summer Session, Chapter II, Part 2, September 8-19, 1952. 41. Gunther, F. C, "Photographic Study of Surface-Boiling Heat Transfer to Water with Forced Convection." Trans. ASME, 73, No. 2, (February, 1951), 115.

-170 -42, Bonilla, C. F., Busch, J. S., Stalder, A., Shaikhplahmud, N, S., and Ramachandran, A. "Pool Boiling Heat Transfer with Mercury." Reactor Heat Transfer Conference of 1956. 43, Clark, J. A. Heat Transfer with Saturated Boiling. ScDo Thesis, Massachusetts Institute of Technology, Cambridge, Mass,, May, 1953. 44. Chang, Y, and Snyder, N. W. 'Heat Transfer in Saturated Boiling," Preprint 104 presented at Third National Heat Transfer Conference, ASME-AIChE, August, 1959. 45. Griffith4, P. "The Correlation of Nucleate Boiling Burnout Data," ASME, Paper No. 57-HT-21, Presented at ASME-AIChE Heat Transfer Conference, August, 1957. 46. Handbook of Chemistry and Physics, 38th Edition, Chemical Rubber Publishing Co., Cleveland, Ohio, 1956. 47. Levy, S. "Generalized Correlation of Boiling Heat Transfer," Trans~. ASME, 81, Series C, No. 1, (February, 1959), 37. 48. Forster, K. and Greif, R. "Heat Transfer to a Boiling LiquidMechanism and Correlations." Ibid, 43. 49. Hirano, F. and Nishikawa, K. "The Phenomena of Boiling - Part lt' Kikai No Kenkyu (Machinery Research), Japan, 8, No. 3, (1956), 317. 50. Hirano, F. and Nishikawa, K. "Boiling of Hard Water.'" Trans. Soc. Mech, Eng., Japan, 19, No. 88, (1953), 33. 51. Hirano, F. and Nishikawa, K. "The Phenomena of Boiling - Part 2',' Machinery Research, Japan, 8, No. 4, (1956), 431. 52, Perkins, A. S and Westwater, J. W. t'Measurements of Bubbles Formed in Boiling Methanol," AIChE Journal, 2, No. 4, December, 1956), 471. 53. Gaertner, R. F. and Westwater, J. W. "Populations of Active Sites in Nucleate Boiling Heat Transfer." Preprint 105 presented at Third National heat Transfer Conference, ASME-AIChE, August, 1959. 54. Hirano, F. and Nishikawa, K. "The Phenomena of Boiling - Part 3,'" Machinery Research, Japan, 8, No. 5, (1956), 521. 55. Treshchov, G. G. "An Experimental Investigation of Heat Exchange Mechanism with Surface Boiling Water." Teploenergetika, 4, Noo 5, (May, 1957), 44.

-171 -56. Nishikawa, K. and Urakawa, K. "Experiment of Nucleate Boiling Under Reduced Pressure." Trans.-Soc. Mech. Eng., Japan, 23, No, 136, (December, 1957), 935. 57. DeBortoli, R, A,, Green, S. J., et al. "Forced-Convection Heat Transfer Burnout Studies for Water in Rectangular Channels and Round Tubes at Pressures Above 500 psia." AEC Research and Development Report WAPD-188, Westinghouse Atomic Power Division, Bettis, June, 1958, 58, Averin, E. Ko The Effect of the Material and of the Mechanical Treatment of the Surface on the Heat Exchange in the Boiling of Water." Izv. Akad. Nauk SSSR, Otd. Tekh. Nauk, No. 3, (1954), 116. 59. Zysina-Molozhen, L. M.."Some Data on the Number of Centers of Vaporization in Boiling on Industrial Heating Surfaces." Problems of Heat Transfer During a Change of State: A Collection of Articles, Edited by SO S. Kutateladze, Publication of State Power"-Press, MoscowLeningrad, AEC-tr-3405, (1953), 155. 60. Griffith, P. "gBubble Growth Rates in Boiling." Technical Report No. 8, Mass, Inst. of Tech,,o Div. of Ind. Corp., Cambridge, Mass, June, 1956, 61. Donaldg M, B. and Haslam, F. "The Mechanism of the Transition from Nucleate to Film Boiling." Chem. Eng. Sci,, 8, (1958), 287. 62. Gunther, F. C, and Kreith, F, "Photographic-Study of Bubble Formation in Heat Transfer to Subcooled Water;" Heat!Transfer and Fluid Mechanics Institute, Preprints of Papers, Stanford University Press, Stanford, Calif., (194-9), 1130 63. Farber, E. A. and Scorah, R, L, "THeat Transfer to Water Boiling Under Pressure."' Trans. ASME, 70, (1948), 3690 64, Daily, J, W. and Johnson, V. E,, Jr. "Turbulence and Boundary Layer Effects on Cavitation Inception from Gas Nuclpi." ASME, Paper No. 55-A-142, November, 1955. Also, Technical Re.ort No. 21, Hydrodynamics Laboratory, M.I.T., July,:1955.

UNIVERSITY OF MICHIGAN iii3111 i ii0 1111111111111111111 111 3 9015 03483 7321