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Proceedings of IMECE2008 2008 ASME International Mechanical Engineering Congress and Exposition

October 31–November 6, 2008, Boston, Massachusetts, USA

IMECE2008-68478

EXPERIMENTAL STUDY OF A TWO-PHASE HEAT TRANSPORT DEVICE DRIVEN BY ELECTROHYDRODYNAMIC CONDUCTION PUMPING

Matthew R. Pearson

Department of Mechanical, Materials, and Aerospace Engineering,

Illinois Institute of Technology Chicago, Illinois, USA

pearmat@iit.edu

Jamal Seyed-Yagoobi

Department of Mechanical, Materials, and Aerospace Engineering,

Illinois Institute of Technology Chicago, Illinois, USA

yagoobi@iit.edu

Proceedings of IMECE2008 2008 ASME International Mechanical Engineering Congress and Exposition

October 31-November 6, 2008, Boston, Massachusetts, USA

IMECE2008-68478

ABSTRACT Heat pipes are well-known as simple and effective heat

transport devices, utilizing two-phase flow and the capillary

phenomenon to remove heat. However, the generation of

capillary pressure requires a wicking structure and the overall

heat transport capacity of the heat pipe is generally limited by

the amount of capillary pressure generation that the wicking

structure can achieve. Therefore, to increase the heat transport

capacity, the capillary phenomenon must be either augmented

or replaced by some other pumping technique. Electro-

hydrodynamic (EHD) conduction pumping has been

demonstrated as an effective method for pumping liquid films

by using DC electric fields and a dielectric working fluid.

Beyond increased pumping capacity, EHD conduction pumping

offers other advantages over capillary pumping, such as active

control of the pumping capacity via the intensity of the applied

electric field. This experimental study demonstrates the

prospects of a macro-scale two-phase heat transport device that

is driven by EHD conduction pumping. Various liquid film

thicknesses are considered. In each case, the performance of the

EHD-driven heat transport device at various electric field

intensities is compared to the capabilities of the same device

under gravity alone. The effect of tilt on the device is also

considered.

INTRODUCTION Heat pipes are well-known for their ability to transport

large amounts of heat over relatively long distances. These

devices use two-phase flow and capillary forces to create a self-

circulating two-phase flow that carries the heat from a heat

source to a heat sink. Liquid is drawn from the condenser to the

evaporator through some wicking structure, whereupon it

evaporates. More information about traditional heat pipes can

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be found in Ref. [1]. Limitations do exist that serve to limit the

heat transport capacity of heat pipes. Frequently, the capillary

limitation (the limited pumping generation that the wicking

structure can provide) is the most limiting factor in heat pipe

operation. For this reason, there has been interest in how the

wicking structure of a heat pipe might be augmented or

replaced by some other pumping method. EHD pumping

mechanisms are one such method.

EHD phenomena involve the interaction of electric fields

and flow fields in a dielectric fluid medium. This interaction

can induce fluid motion by an electric body force. The electric

body force density acting on the molecules can be expressed as

[2]

�� � ��� � ���� �

�� ���

������ ��. (1)

The first term represents the Coulomb force, which is the force

acting on the free charges in an electric field. The second

(dielectrophoretic) and third (electrostriction) terms represent

the polarization force acting on polarized charges. The third

term is relevant only for compressible fluids. EHD pumping

has shown extensive potential due to its simple, lightweight,

non-mechanical design, low power consumption, low

acoustical noise, and the ease with which pumping can be

controlled by adjusting the applied voltage.

A variety of EHD pumping mechanisms are based on the

Coulomb force: conduction pumping, induction pumping, and

ion-drag pumping. All of these methods work by creating

regions of non-zero charge density within the working fluid,

but they differ in the manner of generating the Coulomb force.

Ion-drag pumping relies on the injection of ions into the liquid

from sharp liquid/solid interfaces, while induction pumping

uses an AC travelling wave to attract and repel charge induced

in the liquid due to gradients or discontinuities of electric

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conductivity, often due to temperature gradients [3].

Conduction pumping produces a non-zero charge density in

“heterocharge layers” near the electrodes through bulk electric

conduction through the liquid. As a result, the problems of ion-

drag pumping (specifically, degradation of the electrical

properties of the working fluid and potentially hazardous

operation) and induction pumping (specifically, the need for a

gradient in electric conductivity) can be avoided while still

maintaining the ability to pump using the Coulomb force [3].

There has been some research on the use of EHD

enhancement of a two-phase system such as a heat pipe or

capillary-pumped loop. None of these studies has considered

the use of conduction pumping of a two-phase, stratified, liquid

film. Jones [4] proposed replacing the capillary wick structure

of a heat pipe from the condenser to the evaporator with an

EHD pump that utilized polarization forces to generate

pumping. Jones and Perry [5] demonstrated this concept

successfully, but the performance was poor compared to

existing capillary driven heat pipes, due to a significant

mismatch of the circumferential capillary groove and EHD

pumping capabilities. Loehrke and Debs [6] further improved

the EHD heat pipe of Jones and Perry [5] and were able to

achieve equivalent thermal throughput of conventional axial-

groove heat pipes at an adverse tilt of 1.7 cm compared to only

0.3 cm for the conventional heat pipe. In a later study, Bologa

and Savin [7] used the dielectrophoretic force to enhance the

heat transport capacity in an experimental heat pipe operating

as a two phase thermo-siphon. By enhancing the rate of

condensation with EHD, the heat transport capacity was

increased 53% at an applied voltage of 36 kV.

Enhancement of the heat pipe transport capacity utilizing

the Coulomb force was investigated by Babin et al. [8]. They

used an ion-drag pump to generate the Coulomb force and to

increase the capillary limit of the heat pipe. Using R-11 as the

working fluid and a two-stage ion-drag pump located in the

liquid passage, a 20% increase in the transport capacity was

achieved at an applied voltage of 20 kV. Enhancement of the

heat transport capacity of a monogroove heat pipe with EHD

pumping was investigated by Bryan and Seyed-Yagoobi [9].

The EHD pump was located on the liquid channel in the

adiabatic section of the heat pipe, and the working fluid was

refrigerant R-123. The two experimental goals were to

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determine the magnitude of heat transport enhancement that

could be achieved using the EHD conduction pump and to

demonstrate the controllability and recovery of the heat pipe

during dry out. Both were successfully accomplished. Over

100% enhancement in the transport capacity was achieved

using the EHD conduction pump operating at 20 kV, but at

very low current levels on the order of ten micro-amps. The

EHD pump was also able to provide immediate recovery from

dry out when the heat pipe had been experiencing progressive

evaporator dry out.

At the time of the study by Bryan and Seyed-Yagoobi [9],

the conduction pumping phenomenon was not understood and

the flow generation was incorrectly attributed to polarization

forces (dielectrophoretic force). Once the conduction pumping

phenomenon had been clarified [10], the EHD-driven heat pipe

was revisited by Jeong and Seyed-Yagoobi [11] using

electrodes better optimized for conduction pumping. With these

improved electrodes, even more significant increases in heat

transport capacity were accomplished. For example, under

3 mm of favorable tilt, the total heat transport capacity

increased from 520 W with no EHD to 920 W with an applied

voltage of 10 kV. The improvements were even more

spectacular when a 3 mm adverse tilt was applied to the heat

pipe — heat transport capacity increased from 200 W with no

EHD to 800 W with an applied voltage of 10 kV.

EHD conduction pumping has also been investigated

recently as the pumping mechanism for a two-phase loop. In

studies by Jeong and Didion [12,13], up to 13.2 kPa of

pumping head generation was achieved using an EHD

conduction pump installed in the liquid line, with higher head

generation attainable by simply adding additional electrode

pairs to the pump. Under such conditions, the heat pipe was

shown to provide thermal control capacity for 35.8 W/cm2 of

heat flux.

The present study differs greatly from previous studies of

EHD-conduction-driven heat pipes because it uses the

conduction pumping of a stratified two-phase flow (liquid

film), whereas previous conduction-driven heat pipes have used

a single-phase conduction pump installed in the liquid line.

Conduction pumping has been successfully applied to pump

liquid films using electrodes embedded in the bottom channel

surface [14,15,16]. The pumping of a liquid film has several

Fig. 1. Section views of experimental two-phase heat transport device

Evaporator (51 mm)

Adiabatic/pumping sect ion (152 mm) Condenser (203 mm)

Water inlet Water outlet

Electrode Board

Heater

Glass viewing windows Pressure p ort

Vapor thermocouple port

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advantages, in particular the simplicity of the device because

the channel cross-section can be a simple rectangle without

need for any separation of the liquid and vapor phases for

pumping purposes.

EXPERIMENTAL SETUP The experimental two-phase heat transport device is shown

in Figs. 1 and 2. The total length of the channel is

This length is divided into an evaporator section

adiabatic pumping section (152 mm), and condenser section

(203 mm). The approximate width of the channel is

Glass windows running the entire length of the channel are

situated on the front and top walls of the channel to provide

clear visualization of the device’s internal workings. The

refrigerant HCFC-123 is used as the working fluid.

plates on the bottom of the channel act as the heat transfer

medium in the evaporator and condenser sections. Electrodes

are installed along the bottom of the channel in the adiabatic

pumping section, between the two copper plates of the

evaporator and the condenser.

The electrodes are manufactured as a printed circuit board

(PCB) using etching techniques, with the dimensions shown in

Fig. 3. The board material is FR-4 epoxy glass and the

electrodes are copper with a tin/lead reflow finish. The upper

surface of the PCB contains the electrodes and the lower

surface contains two bus lines (high voltage and ground).

0.4 mm vias provide the electrical connections between the

electrodes and the bus lines. The high voltage supply

delivered by a EW50R12 high voltage power supply by

Glassman High Voltage, Inc. Two feedthroughs

high voltage and one for ground), manufactured by CeramTec

North America, provide electrical connectivity from the high

voltage power supply to inside the pressure-sealed two

channel. These feedthroughs have a voltage rating of

and a current rating of 55 A. For the studies of EHD, a voltage

Fig. 2. Photograph of the two-phase heat transfer device. The

valves and pipes at the bottom-right of the image are for the

connection of refrigerant supply tank, refrigerant recovery device,

and vacuum pump

of the device because

section can be a simple rectangle without

need for any separation of the liquid and vapor phases for

phase heat transport device is shown

. The total length of the channel is 406 mm.

This length is divided into an evaporator section (51 mm),

, and condenser section

the channel is 57 mm.

Glass windows running the entire length of the channel are

situated on the front and top walls of the channel to provide

of the device’s internal workings. The

123 is used as the working fluid. Copper

plates on the bottom of the channel act as the heat transfer

medium in the evaporator and condenser sections. Electrodes

are installed along the bottom of the channel in the adiabatic

pumping section, between the two copper plates of the

are manufactured as a printed circuit board

, with the dimensions shown in

4 epoxy glass and the

electrodes are copper with a tin/lead reflow finish. The upper

surface of the PCB contains the electrodes and the lower

s (high voltage and ground).

vias provide the electrical connections between the

electrodes and the bus lines. The high voltage supply is

high voltage power supply by

Glassman High Voltage, Inc. Two feedthroughs (one for the

high voltage and one for ground), manufactured by CeramTec

North America, provide electrical connectivity from the high

sealed two-phase

feedthroughs have a voltage rating of 20 kVDC

For the studies of EHD, a voltage

fer device. The

right of the image are for the

connection of refrigerant supply tank, refrigerant recovery device,

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of 5 kV was used in all tests. This represented the highest

voltage that could be used without the risk of arcing.

gradients in temperature (and therefore electrical conductivity)

do exist within the liquid film, the DC electric field imposed by

the power supply will produce only conduction pumping, as

induction pumping requires a AC traveling

field. [3] Ion-injection is also assumed t

there is likely to be some minor injection of ions at the corners

of the electrodes.

Heat is delivered to the evaporator by an ULTRAMIC®

600 ceramic heater by Watlow Electric Manufacturing

Company. The nominal power rating of the heate

and its dimensions are 35 mm ×

3 mm. This corresponds to a maximum heat flux of 100

Shin-Etsu X23-7762 thermal compound

improved thermal contact between the heater and t

surface of the evaporator plate. A T

bonded to the lower surface of heater, is used to monitor the

heater temperature. The heater power is controlled by a variable

transformer, with a Watlow 965 temperature controller u

shut off power to the heater if the temperature of the heater

exceeds 110°C. The voltage drop across the heater and the

current flow through the heater are both measured separately

using digital multimeters. The lower surface

insulated using 0.5 in. thick cork to ensure that most of the heat

is delivered through the copper to the refrigerant.

removed at the condenser using a chilled water supply provided

by a NESLAB HX-150 chiller by

Inc.

A pressure transducer and backup pressure gauge are used

to monitor the pressure of the chamber. A pressure release

valve is set to approximately 100

dangerous pressure rises caused by the heating of the working

fluid. A DV-6 vacuum tube from Teledyne Hastings

Instruments is used to ensure a deep vacuum of at least

500 µm Hg exists before charging the device with refrigerant.

Two T-type thermocouple probes are inserted into the chamber,

one in the vapor phase and one in the liquid p

Unfortunately, the liquid temperature reading is greatly

influenced by the temperature of the chilled water, due to the

placement of the thermocouple. This placement was

unavoidable due to the need to handle very thin liquid films.

Therefore, the liquid temperature is not monitored during

experiments and the saturation conditions and refrigerant purity

are judged based on the vapor temperature reading.

Fig. 3. Electrode design (all dimensions in

5.1

5.1

1.22 4.06

152

kV was used in all tests. This represented the highest

voltage that could be used without the risk of arcing. Although

gradients in temperature (and therefore electrical conductivity)

exist within the liquid film, the DC electric field imposed by

the power supply will produce only conduction pumping, as

induction pumping requires a AC traveling-wave-type electric

injection is also assumed to be small, although

there is likely to be some minor injection of ions at the corners

Heat is delivered to the evaporator by an ULTRAMIC®

Watlow Electric Manufacturing

. The nominal power rating of the heater is 1225 W

× 35 mm with a thickness of

This corresponds to a maximum heat flux of 100 W/cm2.

thermal compound is used to provide

between the heater and the copper

. A T-type surface thermocouple,

bonded to the lower surface of heater, is used to monitor the

The heater power is controlled by a variable

transformer, with a Watlow 965 temperature controller used to

shut off power to the heater if the temperature of the heater

C. The voltage drop across the heater and the

current flow through the heater are both measured separately

The lower surface of the heater is

in. thick cork to ensure that most of the heat

is delivered through the copper to the refrigerant. Heat is

removed at the condenser using a chilled water supply provided

chiller by Thermo Fisher Scientific,

A pressure transducer and backup pressure gauge are used

to monitor the pressure of the chamber. A pressure release

valve is set to approximately 100 kPa gage to prevent any

dangerous pressure rises caused by the heating of the working

m tube from Teledyne Hastings

Instruments is used to ensure a deep vacuum of at least

Hg exists before charging the device with refrigerant.

type thermocouple probes are inserted into the chamber,

one in the vapor phase and one in the liquid phase.

Unfortunately, the liquid temperature reading is greatly

influenced by the temperature of the chilled water, due to the

placement of the thermocouple. This placement was

unavoidable due to the need to handle very thin liquid films.

uid temperature is not monitored during

experiments and the saturation conditions and refrigerant purity

are judged based on the vapor temperature reading.

. Electrode design (all dimensions in mm)

16.51

0.97

50.8

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All data presented in this work were obtained at a

refrigerant saturation temperature of 30.0°C ± 0.2°C. The

corresponding saturation pressure, based on saturation tables

for HCFC-123, was 109.5 kPa absolute. If the measured

pressure in the system exceeded 111.5 kPa (due to the leakage

of air into the device during periods of inactivity) then the

refrigerant was recovered, a deep vacuum was pulled, and the

device was charged with fresh refrigerant. For each data point,

the chiller set point temperature was fixed. The heater power

was then manually adjusted using the variable transformer until

the device reached steady state within the chosen saturation

temperature range. Once steady state was reached, all data

readings were recorded. The chiller set point temperature was

then reduced in preparation for the next data point. Chiller set

point temperatures from 30°C to 5°C were chosen in 1, 2 or

3°C increments.

RESULTS Three different film thicknesses were considered: 2 mm,

4 mm, and 6 mm. In all three cases, the heat transport device

was carefully oriented to the horizontal position. The resulting

boiling curves are shown in Figs. 4–6. For all heat fluxes

considered, pool boiling was present in the evaporator.

Although pool boiling is generally avoided in traditional heat

pipes in order to avoid disrupting the capillary pressure

generation, pool boiling provides no problems for the EHD-

driven pumping mechanism.

For the 2 mm film, Fig. 4 shows that without EHD the

burnout of the evaporator occurred at a heat flux of

14.4 W/cm2. Without any EHD pumping, the liquid film flows

due to gravity, with the phase-change processes in the

evaporator and condenser causing a slight gradient in film

thickness and a resulting gravity-driven flow to the evaporator.

From visual observation of the evaporator section, it was clear

that the burnout was caused by dryout of the evaporator, with

insufficient liquid being pumped into the evaporator by gravity.

Application of 5 kV EHD to supplement the gravitational body

force caused immediate re-wetting of the evaporator due to the

enhanced pumping. With EHD active, dryout of the evaporator

did not occur until a heat flux of 51.4 W/cm2, representing an

improvement of 350%.

For the 4 mm case, the thicker film allowed a higher

gravity-driven flow rate to exist. As a result, the no-EHD

burnout heat flux increased to 45.6 W/cm2, as shown in Fig. 5.

In this case, burnout was once again caused by the drying out

of the evaporator due to insufficient liquid flow rate.

Application of 5 kV EHD caused a slight increase in the slope

of the boiling curve, representing an improvement in the heat

transfer coefficient at the evaporator. The burnout heat flux also

increased to a value of 54.8 W/cm2. Furthermore, when the

burnout heat flux was reached with EHD active, liquid was still

present in the evaporator section. Therefore, the mechanism of

burnout is not caused by insufficient liquid flow rate but by

hydrodynamic phenomena occurring very near the evaporator

surface, akin to the critical heat flux (CHF) condition in pool

boiling.

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Fig. 4. Boiling curve for the 2 mm film, no tilt.

Fig. 5. Boiling curve for the 4 mm film, no tilt.

Fig. 6. Boiling curve for the 6 mm film, no tilt..

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Further increasing the film thickness to 6 mm, EHD is

shown to provide very small increases in the slope of the

boiling curve. However, the curves eventually cross such that

the burnout heat flux actually appears to be slightly lower with

the EHD pumping than without it. The burnout condition (a

CHF condition as the evaporator section was still filled with

liquid) occurred at 51.3 W/cm2. Without EHD, this same heat

flux did not cause burnout. Unfortunately, higher heat flux

levels could not be applied because no additional heat could be

removed from the condenser, its performance somewhat

reduced by the thick level of liquid present there.

To further consider the effectiveness of the EHD film

pumping performance, the heat pipe, containing a 4 mm-thick

liquid film, was inclined such that the evaporator side of the

setup was 14 mm higher than the condenser side of the setup.

This inclination was such that without EHD the evaporator was

completely dry, with the “shore-line” of the liquid film aligned

just below the leading edge of the evaporator plate. Because the

evaporator was initially at dryout, no 0 kV data could be

obtained (small heater heat fluxes caused the heater

temperature to reach the maximum allowable temperature). As

such, the burnout condition is artificially marked at 0 W/cm2 in

Fig. 7. However, with 5 kV of applied voltage to the electrodes,

the fluid was pumped very effectively up the incline such that

fluid wet the evaporator. As a result, evaporator dryout and the

associated heater burnout occurred at 20.5 W/cm2. For this case

of adverse tilt, a 2 mm film thickness was attempted, but the

EHD pumping, while present, was not sufficient to wet the

evaporator. For the 6 mm film thickness, so much liquid was

collected in the condenser section due to the adverse tilt that

heat removal by the condenser was extremely limited and heat

fluxes of interest could not be studied.

As an additional study, a 9 mm favorable tilt (i.e. the

evaporator being below the condenser) was considered for the

case of a 2 mm film thickness. Figure 8 shows reasonable

increases in the slope of the boiling curve, suggesting that the

EHD pumping force remained important even in the presence

of a large gravitational body force. Without EHD, the critical

heat flux condition occurred at 51.7 W/cm2. With EHD, heat

fluxes of 56.8 W/cm2 were achieved without burnout

occurring. Further increases in heat flux were not possible due

to limitations in heat removal by the condenser. Thicker films

were not considered for the favorable tilt because of extensive

pooling of liquid in the evaporator section.

For all test cases with an 5 kV applied voltage, the steady-

state current and power consumption of the EHD conduction

pump remained below 100 µA and 0.5 W, respectively, which

is a trivial amount relative to the power input of the heater (e.g.

613 W of heater power corresponds to 50 W/cm2 of heat flux).

DISCUSSION OF RESULTS As a feasibility study of the concept of an EHD

conduction-driven heat pipe, the results are promising.

Conduction pumping head generation has been shown, by

Atten and Seyed-Yagoobi [10] to increase approximately

linearly with the number of electrode pairs, provided that those

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pairs are spaced a sufficient distance apart that they do not

interact. Power consumption would then increase linearly too.

Therefore, the liquid film can be pumped over much longer

distances by extending the adiabatic/pumping section and

incorporating additional electrodes, while retaining very

modest power consumption levels.

From visual observation of the heat transport device during

experiment runs, it is clear that the incorporation of electrodes

in the condenser section would likely serve to significantly

improve performance. In the current configuration, the

condenser represents a very large load on the EHD pump and

even when EHD is enabled, liquid is supplied to the electrodes

in the adiabatic/pumping section due to gravity alone.

Therefore, the film thickness in the condenser decreases

slightly as it approaches the EHD pump, such that for thin films

(e.g. 2 mm), the thickness of the film at this boundary between

the condenser and the adiabatic/pumping section becomes very

small, choking the flow and limiting the EHD pumping

performance. The inclusion of electrodes along the length of

Fig. 7. Boiling curve for the 4 mm film, 14 mm of adverse tilt.

Fig. 8. Boiling curve for the 2 mm film, 9 mm of favorable tilt.

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the condenser would likely eliminate these problems, and the

flow of liquid in the condenser could be significantly

augmented by the presence of an EHD body force. This effect

is evidenced in Fig. 8 where the favorable tilt served to assist

the flow of liquid out of the condenser section and assist the

EHD pumping mechanism. Even with the gravity-driven flow

being very effective, the data show further improvements to the

heat transfer coefficient and heat flux capacity once EHD-

assistance was provided, aided by the gravity-driven flow in the

condenser section. For thicker films, a larger gravitational body

force can exist in the condenser, but the extension of the

pumping section into the condenser is still expected to cause

significant improvements in pumping.

Visual observations also lead the investigators to believe

that the placement of electrodes in the evaporator would also be

extremely beneficial, particularly for thin films and adverse

tilts. For example, Fig. 7 shows the performance of the heat

transport device under a significant adverse tilt. It must be

noted that at such a tilt, electrodes were effective in wetting

only the first 50% of the evaporator length. Consequently, the

area over which evaporation was occurring was approximately

halved. Once the liquid passed the last electrode, it quickly lost

momentum as it travelled uphill along the evaporator. The

presence of additional electrodes to continue the pumping of

liquid along the entire length is expected to yield significant

performance improvements for these adverse tilts by allowing

100% of the area of the evaporator to be wetted.

The electrodes in the current setup do not extend across the

entire width of the channel, as shown in Fig. 3. As a result,

there was generally some backflow of liquid along the sides of

the channel from the evaporator back to the condenser. This

flow could be seen by the trajectory of bubbles departing from

the leading edge of the evaporator – those in the central portion

of the channel departed away from the adiabatic/pumping

section while those very close to the walls of the channel

departed towards the adiabatic/pumping section. Therefore,

removing this gap is expected to result in improved pumping

and heat flux performance due to the higher volume of liquid

film that is subject to an EHD body force, the reduction or

removal of the backflow and associated pumping losses that

occur as a result.

In many aspects, the horizontal, EHD-driven flow of the

current heat transport device is analogous to studies of the

boiling of vertical, falling films that are gravity-driven (see,

e.g., Refs. [17,18,19]). However, the current research has

illustrated notable differences between the two flows. In the

case of vertical flows, studies have considered the flow over a

heated surface but the flowing film is allowed to continue

beyond that surface. Therefore, if the liquid film mass flow rate

entering the evaporator exceeds the mass rate of evaporation

then excess liquid leaves the evaporator. In the current EHD-

driven experimental device, the end of the device is closed;

therefore, any excess liquid mass accumulates in the

evaporator, thickening the film there, until the pressure head of

the thickened film balances the pressure generation of the EHD

pump. From observations of vertical falling films, for example

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by Mudawwar et al. [17], the critical heat flux occurs when the

main part of the liquid film separates from the heated wall due

to the vigorous vapor generation at the surface. In the case of a

vertical film, gravity can do nothing to prevent this separation

as there is no component of gravitational force directed toward

the wall. However, in the case of the horizontal film in the

present study, gravity assists in bringing the liquid to the heater

surface (as it does with pool boiling), which could potentially

lead to higher heat fluxes than for vertical, falling films.

In the present study, no attempt has been made to tune the

voltage applied to the electrodes. It has been mentioned that if

the rate of pumping exceeds the mass rate of evaporation then

some liquid accumulates in the evaporator section, thickening

the film in that location. This thickening may possibly serve to

diminish the maximum heat flux – bubbles can escape from the

surface more quickly when the film is thinner [19]. This

phenomenon may be the reason that the boiling curves in Fig. 6

appear to cross, with the critical heat flux occurring at a lower

heat flux when EHD pumping is active. The thinner film in the

evaporator that exists when EHD is switched off may enable a

higher critical heat flux by allowing the nucleating vapor

bubbles to escape more quickly.

CONCLUSIONS An experimental two-phase device driven by the EHD

conduction pumping phenomenon has been fabricated and

demonstrated. The results have shown that the use of EHD can

provide significant increases in the maximum heat flux of the

device when compared to the use of gravity alone. Performance

improvements have also been demonstrated for both adverse

and favorable tilts. For thin films, the maximum heat flux

corresponds to evaporator dryout, which is caused by

insufficient pumping. For thicker films, the maximum heat flux

corresponds to a critical heat flux condition, as liquid is still

present in the evaporator section. Through examination of the

data and from visual observations of the evaporation and EHD

pumping processes, several improvements to the design have

been proposed. Most notably: (1) The need for electrodes along

the entire length of the device, not just in the central adiabatic

section, in order to provide increased liquid flow rates,

especially for thinner films, and (2) Further study of the boiling

processes in the evaporator and an effort to adapt the models of

vertical, falling films to the case of horizontal film motion.

NOMENCLATURE � = electric field vector

� = electric field magnitude ��� �� = electric body force

� = temperature

= electric permittivity

� = mass density

�� = charge density

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Down

ACKNOWLEDGEMENTS The authors thank NASA Headquarters – Microgravity

Fluid Physics Program and the NASA Goddard Space Flight

Center for their financial support of this research project. The

first author also thanks the National Science Foundation

Graduate Research Fellowship Program for their support of his

graduate studies.

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