boiling heat transfer phenomena from microporous and porous surfaces in saturated fc-72

Pergamon In1 .I Hror Ma.r~ Transfer. Vol. 40, No 18. pp. 44374447. 1997

c 1997 Elsewrr Saence Ltd. All right\\ reserved Prmted m Great Braam

0017-9310~97$1700+0.00

PII : soo17-9310(97)00055-0

Boiling heat transfer phenomena from micro- porous and porous surfaces in saturated

FC-72 J. Y. CHANG and S. M. YOU?

Department of Mechanical and Aerospace Engineering. The University of Texas at Arlington. Arlington, TX 76019-0023, U.S.A.

(Received 31 July 1996 and infinalform 21 January 1997)

Abstract-Particle size effects on boiling performances of micro-porous enhanced surfaces are studied using five different sizes of diamond particles. By comparing the coating thicknesses with the superheated liquid layer thickness, the coatings are classified into two groups : ‘micro-porous’ and ‘porous’ coatings. The superheated liquid layer thickness is calculated using one-dimensional transient thermal conduction. Micro-porous coating shows different characteristics of boiling performances compared to porous coating

in incipient superheat, nucleate boiling and CHF. c; 1997 Elsevier Science Ltd.

1. INTRODUCTION

In recent years, many boiling heat-transfer enhancement studies have been made. One boiling enhancement method of great interest is to increase the number of small-scale cavities on a surface. Techniques for increasing small-cavities include the application of porous coatings, mechanical grooves or dendrites to the desired surface. The micro-geometries of these enhanced surfaces serve to increase vapor/gas entrapment volume and active nucleation site density. These increases combine to reduce incipient superheat and nucleate boiling superheat ( Tw - T,,,) and, hence, increase the boiling heat-transfer coefficient.

Griffith and Wallis [l] demonstrated that the geometry of a cavity containing trapped vapor is directly related to the bubble nucleation process. They determined that the diameter of the cavity mouth defines embryonic bubble size and, hence, the wall superheat at which the cavity will be activated. From their notable studies, re-entrant-type cavities were shown to act as stable, easily activated boiling sites. Several commercial enhanced surfaces have been developed to produce re-entrant-type cavity geometries : ECR-40 of Furukawa Electric, GEWA series of Wieland-Werke, High-Flux of UOP, Thermoexcel series of Hitachi and Turbo-B of Wolverine. Experi- ments have shown that these structured surfaces per- form successfully, as summarized by Thome [2].

While many of the enhanced surfaces tested have demonstrated the ability to reduce wall superheat and increase CHF, their feature sizes are apparently too large to effectively trap a large number of embryonic bubbles when immersed in dielectric liquids (e.g

t Author to whom correspondence should be addressed.

refrigerants) [3]. By manufacturing optimum cavity sizes on a heated surface, both the boiling site density and the nucleate boiling heat transfer can be efficiently increased. You et al. [4] introduced a ‘particle lay- ering’ technique which was applied to a flat surface. Alumina (A1,07) particles (0.3-3 pm) were sprayed on a heater and tested in FC-72. The deposited particles adhered to the surface due to van der Waals molecular attraction forces. They reported significant reductions in incipient and nucleate boiling superheats (Z 50%) and an increase in CHF (~32%). Based upon the concept by You et al. [4], O’Connor and You [5] developed a boiling enhancement paint with silver flakes (3-10 pm). Their treated surface, immersed in saturated FC-72, showed an ~80% reduction in nucleate boiling superheat and a z 109% increase in CHF over the non-painted surface. O’Connor et al. [6] developed a dielectric paint, made with diamond particles (8812 pm), for applications in electronic component cooling. Their dielectric paint surface showed nearly identical pool boiling enhancement with that of the silver flake paint. A painted silicon chip tested at 45 K subcooled condition showed a 224% larger CHF value (159 W cm- ‘) than that of the untreated chip surface at saturated conditions. More recently, Chang and You [7] examined the effects of heater orientation on CHF and nucleate boiling heat transfer for uncoated and coated flat surfaces. They used two different coatings which were made with copper particles (l-50 pm) and aluminum particles (I-20 pm).

The present research is an experimental study of pool boiling heat transfer from diamond-particle- coated surfaces immersed in saturated FC-72. Exper- iments were performed to understand the effects of particle size on pool boiling heat transfer from the enhanced surfaces. Five different sizes of diamond

4437

4438 J. Y. CHANG and S. M. YOU

A CHF CHF,

area [m’] critical heat flux [W cm- ‘1 CHF predicted by Zuber [lo] [W cm-‘]

C, specific heat of heater material [W kg--’ Km’]

d, R h

h, k In n: P Pr

4”

mean particle diameter [m] gravitational acceleration [m s- ‘1 heat-transfer coefficient, q”/( T,, - T,,,) [W crnm2 Km’] heat of vaporization [J kg ‘1 thermal conductivity [W mm ’ K ‘1 mass [g] active nucleation site density pressure [Pa] Prandtl number heat flux [w cm-‘]

r, C,, and s empirical constants in equation (1)

rb embryonic bubble radius [m] T temperature [K] t thickness [m]

NOMENCLATURE

G; velocity of vapor jet [m s ‘1 V volume [ml].

Greek symbols 6 99 superheated liquid layer [m] 1: porosity [%]

*H Taylor instability wavelength

[ml P viscosity [kg mm’ s ‘1

P density [kg mm ‘1 rJ surface tension [N mm’].

Subscripts bulk bulk liquid C coating crit critical f liquid state

g vapor state h heater material

P particle sat saturation state w heater surface.

particle (2 pm-&l, 10 prnS2, 20 pmf3, 45 pm+5 and 70 pm f IO) were used in the experiments.

2. EXPERIMENTAL APPARATUS AND

PROCEDURE

2.1. Test,facilitl The pool boiling test facility is shown in Fig. 1. The

test liquid was contained within a glass vessel and submerged in an isothermal water container. Atmo- spheric pressure was maintained by venting the vessel to ambient. Two copper-constantan thermocouples were placed within the test vessel to measure bulk liquid temperature. The heater assembly was mounted horizontally to an aluminum support assembly and immersed in the test liquid.

Figure 2 shows the test heater assembly. Serpentine windings of a nichrome wire (0.144-mm diameter) were attached to a Teflon substrate (1 l-mm thick) using Omegabond 200 high temperature epoxy (k = 1.4 W m ’ K ‘). A 1.5-mm-thick block of copper (10 x 10 mm) was bonded on top of the heating element using the same epoxy. Two layers of epoxy assured electrical insulation. To provide surface temperature measurements, two copper-constantan thermocouples (30 wire gage, 0.255-mm diameter) were inserted and soldered into two holes (1 -mm diameter and 5-mm depth) drilled in the center of the copper block. A shunt resistor was used to determine the DC current in the electric circuit. Heat flux was estimated using the measured voltage drop across the test heater.

Thermocouple to Condenser

from Condenser

Immersion heater/circulator

Lexan water vessel

Glass test vessel

Fig. I Schematic of test apparatus.

2.2. Test procedure The water bath was heated to the test liquid’s

saturation temperature using the immersion heater/ circulator. Once the test liquid reached its saturation temperature, it was left at this state for 2 h to remove dissolved gases. A magnetic stirrer was used

Boiling heat transfer phenomena 4439

Copper block (10 mm x 10 mm) with two thermocouple holes of 1 -mm diameter and 5-mm depth

Teflon substrate

Nichrome wire insulated with Omegabond 200

Fig. 2. Test heater geometry

during this process to accelerate the dissolved gas removal. After degassing, the magnetic stirrer was turned off and data acquisition begun.

The heat-flux was controlled by voltage input. After each voltage increase (heat-flux increment), a 15 s delay was imposed before initiating data acquisition. After the delay, the computer collected and compared temperature measurements from each thermocouple until the temperature differences for all thermocouples were less than 0.2 K. The test section at this point was assumed to be steady-state. A steady-state condition was usually reached approximately 45 s after each heat-flux increment. After reaching steady-state, the bulk fluid temperature was measured and the heat- flux was calculated.

For heat-flux values greater than ~80% of CHF, the instantaneous surface temperature was monitored for 45 s after each increment to prevent heater burnout. Each instantaneous surface temperature measurement was compared with the previous increment’s average surface temperature. If a temperature difference larger than 30 K was detected, CHF was assumed and the power was shut down. The CHF value was computed as the steady-state heat- flux value just prior to the power supply shutdown plus half of the increment. The heat-flux increment used near CHF was ~0.5 W cmm2. At least two con- secutive runs were conducted for each surface tested. Details of the test equipment and the test procedure were described by Chang and You [7].

Substrate conduction losses were estimated based upon numerical analysis conducted by O’Connor and You [5]. Heat losses were estimated to be between I5 and 5% for heat-fluxes between 0.5 and 15 W cm-‘, respectively. The overall uncertainty estimates in heat- flux were 15.5 and 5% for these heat-flux conditions. The uncertainty values of the present heater can be estimated to be smaller than those of O’Connor and You for the larger surface area (1 .O vs 0.825 cm’) with comparable thickness. Also, temperature measure-

ment uncertainties were estimated considering thermocouple calibration errors, temperature correction for the embedded thermocouples and thermocouple resolution errors. The uncertainties for measured average surface superheat were k 0.5 and k 0.4 K for the heat-flux conditions referred to above. Based upon the uncertainties in the heat-flux and the surface superheat, the uncertainty estimates in the heat transfer coefficient were 16 and 5% for the heat-flux conditions referred to above.

3. RESULTS AND DISCUSSION

All experiments were conducted in pool boiling of saturated FC-72 at 1 atm. The tested surfaces included a smooth ‘plain’ reference surface and plain surfaces coated with five different coatings of different particle sizes.

3. I. Pool boiling tests ofplain surfaces The plain surface was tested first. Heater thickness

effects on pool boiling cures were investigated. Cop- per-block heaters with three different thicknesses (f,, = 3.2, 1.5 and 0.58 mm) were compared. The heater design used for the heater with the 1.5-mm- thick copper block (shown in Fig. 2) was also used for the heater with the 3.2-mm-thick copper block. In the case of the 0.58-mm copper block, three smaller thermocouples (40 wire gage, 0.0799-mm diameter) were directly soldered to the bottom surface of the copper block and spaced evenly across the centerline. The thermocouples were located between the heating element and the copper block. The copper surfaces of the heaters were polished using Brasso. After polish- ing, a small amount of 3M epoxy (1832L-B/A) was carefully applied around the perimeter of the heater surface to prevent undesired edge nucleation sites. Figure 3 illustrates the representative pool boiling curves for these plain-surface heaters with three different heater thicknesses. Negligible differences in

4440 .I. Y. CHANG and S. M. YOU

1.-i I l#ll!llllll I III , I ,1,,1/,,,1,,,1 I I II

O0 2 3 4 56 10’ 2 3 4 56 lo2

Surface superheat ( K ) Fig. 3. Heater thickness effects on pool boiling curve.

4

3

10-l

natural-convection data indicate small substrate conduction losses for the present heater design. The heater surfaces produced nearly identical nucleate boiling curves independent of the heater thicknesses.

The effects of the heater thermal properties and the heater thickness on CHF were investigated by Bar- Cohen and McNeil [8] and Carvalho and Bergles [9]. They collected available CHF data and tabulated heater thermal properties and thicknesses. These CHF values were correlated with the combined variable, I,_/=, which was termed the thermal activity parameter (conductance and capacitance). Zuber’s [IO] correlation predicts CHF, = 15.1 W cme2 for saturated FC-72. The CHF data produced from the plain surfaces with three different thicknesses were then normalized based on CHF,. The normalized CHF values vs the thermal activity parameter are plotted in Fig. 4. The present CHF data show a negligible heater thickness effect for the three tested thicknesses. Most of the present CHF data are within k 10% of Zuber.

3.2. Pool boiling tests of micro-porous-enhanced surfaces

Particle size effects were studied using diamond particles, which possess the benefits of uniform size dis- tributions and a wide range of available sizes. The diamond particles were used to make a composition identified as ‘DOA’ coating, so named because each letter stands for a component of the coating : D for diamond particles, 0 for Omegabond 101 epoxy and A for alcohol (isopropyl). In this study, five DOA coatings with different particle sizes were fabricated.

Table 1 shows the particle size of each DOA coating used in the present study. Scanning electron micro- scope (SEM) images of the top views of five DOA coatings (2, 10, 20, 45 and 70 pm) are shown in Fig. 5. Even though a significant difference in particle size is observed, the coatings show a similar particle shape of an edged polygon. During this study, three different plain heaters were used to assume repeatability for each DOA coating. Coated surfaces were generated consistently using the dripping method. Coating thicknesses were measured with the micrometer and confirmed with SEM pictures of cross-sections of the coated samples. The measured thickness of each DOA coating is presented in Table 1. Each coating thickness was maintained at around five times its mean particle diameter, except for the DOA coatings with d,,, = 2 and 70 pm, where the resulting thickness consistency was t,/d,,, z 15 and 3.6, respectively.

3.2.1. Micro-porous andporous coatings. O’Neill et ul. [ 1 l] proposed a nucleate boiling model for a porous coated surface. In their model, a thin superheated liquid film was assumed to exist on the surface of each particle stacked within the porous layer. They proposed that heat is transferred by conduction through the particle matrix and then by conduction across the thin liquid film where evaporation occurs. This model was based upon the higher thermal conductivity value (= 400 W m- ’ K -‘) of sintered copper particles employed in their high-flux surface. However, for the present coatings, O’Connor et al. [6] discussed the coating layer not being a good thermal conductor. The low thermal conductivity value was thought to be due to the epoxy component within

1.0

LLN I 0.8 0

2

5 0.6

0.4

Boiling heat transfer phenomena

. . . . _:. . . . A t = 3.2 mm Cl t = 1.5mm 0 t ~0.58 mm 1

_____________._. :... - Bar-Cohen & McNeil (1992) : .-.- Carvalho & Bergles (1992)

1

0 20 40 60 80 100 120

t ( ph Cph kn )“‘5 ( J/rr~Ksec~‘~ ) Fig. 4. The influence of thermal-activity parameter on CHF.

140

Table I. Particle sizes used for DOA coatings

Particle size Coating (d,) thickness (13

Porosity of powder (E)

2pm+l 30pm*lO =I5 48% 10 pmf2 50 pm*20 -5 41% 20 pm*3 100 pm+ 30 -5 40% 45 pm*5 200 pm * 50 -4.4 40% 70pm&lO 250 pm + 50 -3.6 47%

the coating which impeded conduction between the micron-size metal particles. Based upon this low thermal conductivity, a superheated layer is assumed to develop from the heater base surface through the coating layer.

From the study by Hue [12], a transient thermal boundary layer concept was used over a heated surface to determine activation of cavities during nucleation. Based upon Hue’s model, the superheated liquid layer thickness, &,,, was estimated for the present test liquid (FC-72). A transient conduction solution [ 131 for a constant heat-flux surface in an infinite medium was generated assuming bubble departure frequency was 46 Hz [14]. From this analysis, the superheated liquid layer thickness, a,,, was computed to be approximately 100 pm. This estimation proposed that for the coatings in Table 1 whose thickness (tJ was greater than 100 pm, only the lower portion of the coating layer is activated during nucleate boiling. By comparing coating thicknesses with the estimated &, the DOA coatings in Table 1 are classified

4441

into two groups: ‘micro-porous’ and ‘porous’ coatings. The micro-porous coating thicknesses are less than & (KY 100 ,um) and the porous coating thicknesses are greater than &,. Figure 6(a), (b) shows theoretical models of micro-porous and porous surfaces, respectively. The present porous surface is especially named as a ‘porous nonconducting surface’ by considering the difference from the highly con- ductive coating of O’Neill et al. which is named as a ‘porous conducting surface’. Figure 6(c) shows a theoretical model of the porous conducting surface where a superheated liquid layer develops over the entire porous structure.

In Fig. 7, nucleate boiling performance characteristics (heat-transfer coefficient h vs heat flux) of the micro-porous and the porous nonconducting DOA coatings are compared for heat fluxes below 15 W cm-2. For comparison, nucleate boiling data for an uncoated, plain surface were also plotted. A similar value of slope is observed for all the micro-porous surfaces and decreased values of slope are observed for the porous nonconducting surfaces. These different trends in slope evidently show the difference in nucleate boiling characteristics between the two groups (micro-porous and porous nonconducting surfaces). Considering the micro-porous structures shown in Fig. 6(a), the enhancement of h values for the increasing particle size can be explained by the increased active nucleation sites. As shown in Table 1, a thicker coating was generated with larger size particles. As the coating thickness and pore size are increased, the micro-porous surfaces produce more active nucleation sites as far as the coating thicknesses


Fig. 5. SEM images of DOA-coated surfaces (top views) : (a) 2 pm+ I; (b) 10 pm? 2; (c) 20 pm? 3 ; (d) 45jtm+S;(e)70pmflO.

are smaller than 6,,. For this reason, an optimum coating thickness closer to &,, is recommended for the present micro-porous coatings to utilize the thickness benefit.

The nucleate boiling performances of the porous nonconducting surfaces [Fig. 6(b)] are also compared in Fig. 7. At heat fluxes lower than ~2.5 W cm-*, the porous nonconducting surfaces with larger particle sizes and thicker coatings produced higher heat-transfer coefficients. In this heat-flux range, discrete vapor bubbles were evidently observed to depart from the coating surfaces, rather than larger merged vapor bubbles which were observed above this heat-flux range. These less-crowded vapor bubbles can escape the porous coating layer with little hydraulic resistance. Also, larger sized departing bubbles were observed from larger particle sizes and thicker coatings. Based

upon these observations, the increase of h over the discrete vapor bubble regime was thought to be due to increased active nucleation site density for the larger particle sizes within the extended superheated liquid layer. At heat fluxes higher than ~2.5 W cmm2, this trend was reversed ; lower h was observed with larger particle size coatings. This decrease in nucleate boiling performance can be attributed to the penalty associated with the increased coating thickness by two ways. First, the liquid supply to the innermost active portion (<S,,) of the porous nonconducting coating is hin- dered due to additional hydraulic resistance over liquid-vapor exchange channels, reducing cooling performance. Secondly, the increase of coating thickness produces additional thermal resistance due to the low thermal conductivity of the coating layer, increasing the temperature drop at higher heat fluxes.

Boihng heat transfer phenomena 4443

1 .$$g$g superheated liquid layer

(4

Tsa, Tw T

(W

Fig. 6. Theoretical models of micro-porous and porous coated surfaces : (a) micro-porous surface ; (b) porous non-

conducting surface; (c) porous conducting surface.

2.2

2.0

1.8

1.6

1.4

1.2

1.0

0.8

0.6

0.4

0.2

Scurlock [15] observed coating thickness effect in nucleate boiling heat transfer similar to the present micro-porous and porous nonconducting surfaces. He produced enhanced surfaces using plasma-sprayed aluminum particles (l&100 pm), which were tested within cryogenic liquids (N2, Ar and OJ. He observed enhanced nucleate boiling performance by increasing coating thickness from 130 to 250 pm, as observed with the present micro-porous surfaces. For the tests with six different coating thicknesses from 250 to 1320 pm, he observed the same thickness benefit and penalty over the nucleate boiling performance across a heat flux of z 1.5 W cme2 as observed with the present porous nonconducting surfaces. Assuming his plasma-sprayed surfaces belong to the ‘porous conducting’ group, it is interesting that his observations are favorably comparable with those of the present DOA coatings. As a reference, the superheated liquid layer thickness for liquid nitrogen is estimated to be - 100 pm. For this estimation. the bubble departure diameter was calculated from the correlation proposed by Cole [16], whose result was then used to calculate the bubble departure frequency from the correlation by McFadden and Grassmann [ 171. Con- sidering similar boiling heat-transfer behavior between the present and Scurlock’s surfaces, the increased optimum thickness (- 250 pm) over the theoretical a99 estimation (- 100 pm) from Scurlock’s surface can be due to higher thermal conduction characteristics of the plasma-sprayed aluminum matrix.

The experimental data in Fig. 7 were analyzed in

Open symbols : Micro-porous surfaces Closed symbols : Porous nonconducting surfaces

1 .

- Plain . . ..fIk+ 2p_rn

..D. 10Fm . .O, 20pm

..P. 45 pm . . ..+. 7opnl

Heat flux ( W/cm2 )

Fig. 7. Particle size effects on nucleate boiling enhancement (micro-porous, f, < a,, and porous nonconducting, r, > &, surfaces).


4

3

$2

1

0 L

Pbrous 1 j pqrous ; ngncondycting ; ; 3 \:‘“-i : i i). surfaces . surfaces: ~

61 I I I I I1 1 I I I I I I I

0 10 20 30 40 50 60 70 80

Mean particle diameter ( pm ) Fig. 8. Variations of I/r and C,, with particle size

detail using the theoretical pool boiling correlation by Rohsenow [18]. He proposed a correlation for nucleate pool boiling heat transfer based on his micro- convection model, which takes the form :

C,,(Tw - T,ad hfgPrF = Csf [&&$)J. (l)

Here the coefficient C,, and the exponent Y should be empirically determined to account for liquid-surface combination and the effects of surface preparation. Rohsenow suggested that the exponent s be 1.0 for water and 1.7 for other fluids. Using the least-squares fitting method, the slope (I/r) of the data was estimated and a corresponding C,, value was generated for each surface in Fig. 7. During the curve-fitting procedure, the fitted data agreed with the experimental data within the standard deviations of 7.9, 4.4, 11.7, 17.7, 12.2 and 3.0% for plain, 2, 10, 20, 45 and 70 pm surfaces, respectively. The estimated (l/r) and C,r values are compared in Fig. 8. For simplicity, each surface is represented by a mean particle diameter, d,,,. The uncoated plain surface is represented as a case of zero mean particle diameter. As seen in Fig. 8, the micro-porous surfaces show a similar slope (l/r), which is about 14% higher than that of the plain surface. For the porous nonconducting surfaces, however, as the particle size is increased, the slope is decreased lower than that of the plain surface. As previously discussed in Fig. 7, these different trends were due to the combined effects of additional hydraulic and thermal resistances over the thick porous nonconducting coatings. Carey [ 191 reviewed many different nucleate pool boiling models suggested so far. From his study, it was generally agreed upon

0.020

0.015

0.010 ""

0.005

0.000

that the nucleate boiling heat flux (4”) exhibits a power-law dependence on the active nucleation site density (ni) and surface superheat (Tw- T,,,). Based upon his study, the coefficient Csf in equation (1) is supposed to be affected mainly by the active nucleation site density for a given liquid-surface combination. For the estimated C,, values in Fig. 8, although slightly decreasing and increasing trends are observed across two surface groups (micro-porous and porous nonconducting surfaces), all the coated surfaces show about 70% lower C,, values than that of the plain surface. These even lower C,, values prove that active nucleation site density was effectively enhanced by the present DOA coatings.

3.2.2. Effects qf’particle size on incipience. The variations of incipient superheat for the five

different particle sizes are compared in Fig. 9. Incipi- ent superheat data from three different heaters were combined to represent each particle size. For each heater, more than five runs were made with a time interval of about 1 h between runs. As the particle size is increased over the micro-porous coating regime (0 < d,,, < 20 pm), the incipient superheat decreases significantly. Bar-Cohen and Simon [20] discussed a correlation for incipient wall superheat required to initiate boiling in a highly wetting liquid :

Pv,t(Tw) -P,,,(Td = $ where

u, = 0.042705 I2532

for FC-72.


~

data ____; _______ i _._____ j _______ j . . . . i .._.... i...

1

I

0 10 20 30 40 50 60 70 80

Mean particle diameter ( pm ) Fig. 9. Incipient superheat vs particle diameter.

Using vapor pressure values corresponding to the measured incipient superheats in Fig. 9, the embryonic bubble radii, r,,, were estimated to have ranges of 0.0474.13 pm, 0.14-0.61 pm, 0.47-1.1 pm and 0.79- 2.3 pm for d,,, = 0, 2, 10 and 20 pm, respectively. These estimated embryonic bubble sizes were fairly comparable to the cavity sizes observed from the surfaces in Fig. 5. As the particle size was increased further, similar ranges of incipient superheat (< 10 K) were observed. As previously illustrated by Bankoff [21], vapor/gas entrapment from the advancing liquid film over surface cavities is strongly affected by the liquid contact angle and the cavity geometry. For the porous nonconducting surfaces, it can be assumed that a similar size of cavity geometry was used to capture embryonic bubbles, whereas bigger cavity geometries (&, > N 2 pm) were flooded with the contact of FC-72 liquid film.

3.2.3. Effects of particle size on CHF. CHF variations for five DOA coatings are compared in Fig. 10. The DOA coatings consistently show enhanced CHF values over an uncoated, plain surface. These enhanced CHF values show the CHF dependence on surface micro-structures which was not accounted for by Zuber’s [lo] prediction :

where

CHF, = p&, U, 2 (3) W

u, = JW.

In equation (3), U, is the velocity of vapor jet ; A, is

the vapor jet cross-section area ; A, is the heater surface area ; and 1, is Taylor instability wavelength. Recently, some research studies have been conducted to explain the role of surface micro-structures on CHF. Polezhaev and Kovalev [22] modified Zuber’s CHF theory for porous coated surfaces. They suggested that the spacing of the vapor jets is influenced by the substructure of the porous surface, rather than by the Taylor instability wavelength. Because the dis- tances between the neighboring large pores (dominant vapor exits) on the porous surfaces were observed to be smaller than the Taylor wavelength from the plain surface, the vapor jet velocity was thought to increase substantially from equation (3), resulting in CHF increase. Based upon their model, they generated a CHF equation which included the parameters of the porous coating-porosity and pore size. Katto [23] extended Zuber’s hydrodynamic instability model and related the CHF with Hehnholtz instability of small feeder jets in the macrolayer existing under the unstable vapor blanket. Tehver [24] employed Katto’s CHF model to explain enhancement of CHF by porous coatings. He proposed that macrolayer evaporation time was substantially increased by the pres- ence of the porous layer and, hence, the onset of film boiling was retarded. Using his experimental data with plasma-sprayed aluminum coatings, he proposed lin- ear dependence of CHF on the produce of porosity and mean pore radius.

To compare with the previous models, the porosit- ies and the pore sizes of the present coatings were estimated. For the porosity data, the porosity of the

4446 .I. Y. CHANG and S. M. YOU

30

h 20

NE

2

!k 0

10

0 0 10 20 30 40 50 60 70 80

Mean particle diameter ( pm ) Fig. 10. CHF vs particle diameter.

powder component before mixing with a binder and a carrier was measured using a glass tube graduated with the minimum degree of 0.1 ml and a weighing scale with a readability of 0.001 g. A certain amount of powder volume was put into the glass tube and settled well by tapping the glass tube on the floor. If no change of powder volume was observed, then the volume (V,) and the mass (m,) of the powder was measured. With the measured V, and q,, the porosity of the powder was estimated :

At least five measurements were made for each coating using five different volumes. Test results of the porosity measurements are presented in Table 1, each of which has a nearly identical scatter band of f 2%. For the pore size data, SEM images of the top views of the coated surfaces were used to estimate the pore sizes. From highly magnified views up to 4000 times, a wide range of pore sizes was observed for the present DOA coatings and the large pore sizes were found to be close to the mean particle diameters. The porosity and pore size data estimated from the above procedure were applied to the prediction models of Polezhaev and Kovalev [22] and Tehver [24], and the computed results were compared with the present experimental CHF data. However, the previous models did not give satisfactory explanation for the current experimental CHF data. Since the previous models were developed with porous conducting surfaces, application of these models to the present micro-porous and porous nonconducting surfaces might not be appropriate.

For the present coatings, an empirical approach is used to correlate CHF data with the coating parameters. Considering similar porosity data among the present coatings in Table 1, the CHF data in Fig. 10 is correlated solely with the mean particle diameter, d,,, (z large pore size), by curve fitting :

CHF = 20.62+3.11 In (d,,,-0.25(ln(dm))2

-0.016(ln(d,,,))i. (5)

For the fitted line in Fig. 10, the CHF increase dra- matically at the micro-porous surface regime, and then flattens out at the porous nonconducting surface regime. For the present coating surfaces, the enhance- ments of surface area were not observed by comparing the natural convection portions of the coated surfaces data with that of the uncoated plain surface data (not shown). Accordingly, the increase of CHF by the present coating is due to the enhancement of the active nucleation site density within the coating layer, which effectively delayed the onset of film boiling. This enhancement was evidently observed over the micro- porous surface regime. However, this enhancement mechanism seems to reach a limit and, hence, no more enhanced CHF is observed over the porous nonconducting regime.

4. CONCLUSIONS

Six heater surfaces (plain and five DOA coatings) were tested in a pool of saturated FC-72 at atmo- spheric pressure.


(1) Particle size effects on boiling performance were cation to cooling of electronic equipment. IEEE Trans-

studied using the five different DOA coatings with actions of the CPMT, 1991, 15(5), 90-96.

different sizes of diamond particle (2 prnf 1, 10 5. O’Connor, J. P. and You, S. M., A painting technique

to enhance pool boiling heat transfer in saturated FC- pm+2, 20 p&3, 45 pm&-5 and 70 p&10). For the 72. ASME Journal of Heat Transfer, 1995, 117(2), 387-

nucleate boiling regime, using a transient conduction

impedance for liquid-vapor exchange channels and

solution, a superheated layer thickness, &,, was estimated to be

higher thermal resistance.

z 100 pm for the present flat heater. By comparing the coating thicknesses with the &, the present coatings were classified into two groups. For coatings thinner than a,, (named ‘micro-porous’ coatings), additional thickness due to increased particle size generated a higher active nucleation density. For coatings thicker than 6,, (name ‘porous nonconducting coatings’), two distinct boiling heat-transfer regimes were observed. Below ~2.5 W cm-‘, due to increased active nucleation site density within the extended superheated liquid layer, the coatings with larger size particles showed higher enhancement. About ~2.5 W cm-‘, decreased boiling performance was observed for the coatings with larger size particles due to the thickness penalty associated with higher

the porous nonconducting surfaces, due to a similar size of cavity geometry used to capture embryonic bubbles, similar ranges of incipient superheat (< 10

(2) A significant decrease of incipient superheat was observed over the micro-porous surface regime. For

393. 6. O’Connor, J. P., You, S. M. and Price, D. C., Thermal

management of high power microelectronics via immersion cooling. IEEE Transactions of the CPMT. 1995, N(3), 656-663.

7. Chang. J. Y. and You, S. M., Heater orientation effects on pool boiling of micro-porous-enhanced surfaces in saturated FC-72. ASME Journal of Heat Transftir, 1996, 118(4), 937-943.

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related with the coating parameter [equation (5)]. Sig- 1996, 118, 117- 123.

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between bubble frequency and diameter during nucleate Acknowledgements-This study was supported by the Texas Higher Education Coordinating Board, Advanced Research; Technology Program grant number 003656-014. The authors extend their thanks to the 3M Industrial Chemical Products Division for the donation of FC-72 test liquid.

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boiling heat transfer phenomena from microporous and porous surfaces in saturated fc-72

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