an electrolysis bubble actuated micropump

JOURNAL OF MICROELECTROMECHANICAL SYSTEMS, VOL. 16, NO. 5, OCTOBER 2007 1095

An Electrolysis-Bubble-Actuated MicropumpBased on the Roughness Gradient Design

of Hydrophobic SurfaceChih-Ming Cheng and Cheng-Hsien Liu, Member, IEEE, Member, ASME

Abstract—A novel electrolysis-bubble-actuated micropumpbased on the roughness gradient design in the microchannel isreported in this paper. This micropump is implemented by tak-ing advantage of both the electrolysis actuation and the surfacetension effect. The surface tension effect is controlled via theperiodic generation of electrolytic bubbles and the roughnessgradient design of microchannel surface, which results in thespecified variation of liquid contact angle along the microchannel.Our proposed micropump could resolve the disadvantages thatexist in the early reported micropumps, such as the complicatedtime-sequence power control, the need of long nozzle-diffuserstructure, and the choking/sticking phenomena of electrolytic bub-bles in a microchannel. Due to the features of large actuationforce, low-power consumption, and room temperature operation,our micropump is suitable for the development of low-powerconsumption and compact micropumps for various applications.Experimental results show that the liquid displacement and thepumping rate could be easily and accurately controlled by adjust-ing the amplitude and frequency of the applied voltage. With theapplied voltage of 15 V at 4.5 Hz, a maximum pumping rate of114 nl/min is achieved for one of our micropump designs with amicrochannel of 100 × 20 µm. In this paper, we report the theo-retical analysis, design, micromachining process, operating prin-ciples, characterization, and experimental demonstration of thesemicropumps. [2006-0174]

Index Terms—Bubble, electrolysis, microelectromechanical sys-tems (MEMS) micropump, roughness gradient, surface tension.

I. INTRODUCTION

MANY micropumps based on different designs and ac-tuation mechanisms have been demonstrated over the

past decade [1]. For example, diaphragm pumps actuated bypiezoelectric [2]–[4], electromagnetic [5], thermopneumatic[6], and electrostatic [7] mechanisms have been developed toachieve a high pumping volume through a large chamber. Mostof the fabrications of diaphragm-based pumps are complicatedbecause of the many photolithographic steps involved. Field-driven micropumps such as electro-osmotic [8], [9] and electro-

Manuscript received August 31, 2006; revised April 14, 2007. This work wassupported in part by the National Science Council of Taiwan, R.O.C., underGrant NSC-94-2218-E-007-013 and in part by the Nano-Technology ResearchProgram under Grant NSC-94-2120-M-009-015. Subject Editor A. Ricco.

The authors are with the Micro-Systems and Control Laboratory, Departmentof Power Mechanical Engineering, National Tsing Hua University, Hsinchu30013, Taiwan, R.O.C. (e-mail: [email protected]).

Color versions of one or more of the figures in this paper are available onlineat http://ieeexplore.ieee.org.

Digital Object Identifier 10.1109/JMEMS.2007.900880

hydrodynamic pumps [10] have also been proposed to drive thefluid by applying a high voltage without the need of mechanicalmoving parts. Among all actuation methods, bubble-actuatedvalveless micropumps show the features of simple operation,miniaturized size, large actuation force, and the ability to beconformed physically to different designs of microchannelswith a wide range of cross sections. Both thermal and elec-trolytic bubbles have been demonstrated with the function ofmicrofluidic actuation. The thermal-bubble-actuated microp-umps have been developed with different mechanisms such asthe traversing multiple bubbles [11], the nozzle-diffuser struc-ture [12], and the periodic generation of a single vapor bubblerelying on surface tension [13], [14] to achieve the net pumpingflow. However, the thermal generation of bubbles is a high-power consumption method because of the dramatic heat lossin microscale [12], [15]. The high-temperature process mightalso damage the targeted biosample for biochip applications.

Compared with other actuation mechanisms, the electrolysisbubble actuator has the features of simple structure, largeactuation force, low-power consumption, room temperatureoperation, and being easy to be integrated into a lab chip.The electrolytic bubbles have been used as actuators in var-ious designs of microvalves and micropumps. For example,a microinjector is reported by using the electrolytic bubbles,which raises the liquid pressure to push the liquid forward, andthe bubble is then expelled with the liquid [16]. Micropumpsactuated by electrolytic bubbles are reported and based oninflating a large bubble inside a reservoir, which generates apressure head in the fluid channel during the bubble growth[17], [18]. The electrolytically actuated micropump by meansof sequentially generating a series of electrolytic bubbles insidea microchannel has also been reported [19]. However, there aresome disadvantages for the above pumping mechanisms, suchas the complicated time-sequence power control on many pairsof electrodes, the need of large/long nozzle-diffuser structure,and the requirements of the degassing function and the well-sealed reservoir inside the fluidic chip.

The utilization of surface tension force promises a powerfulactuation mechanism for microfluidic systems because of thelarge force compared to the other forces in microscale. In ourmicropump design, the pumping principle relies on the controlof the surface tension effect that is initiated via electrolysis bub-ble generation and tuned via the roughness gradient design ofhydrophobic surface in the microchannel. Although the similarsurface tension effect is used in our development, the operation

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1096 JOURNAL OF MICROELECTROMECHANICAL SYSTEMS, VOL. 16, NO. 5, OCTOBER 2007

principle and the design in our proposed micropump are differ-ent from those used in other pumping devices [13], [14], [20]–[22]. Our proposed micropump could resolve the challengingproblem for most electrolysis bubble actuators on the removalof insoluble gases. The roughness gradient structure is firstintroduced and utilized to design surface energy gradients toform the guide rails by Shastry et al. [23]. Droplets can movealong the surface energy gradient overcoming the hysteresis bysupplying energy through mechanical vibration. Lee et al. haverecently reported the device that switches the roughness of amembrane to move a droplet [24]. In this paper, we presenta novel electrolysis-bubble-actuated micropump utilizing thedesign of the roughness gradient surface to achieve the pumpingfunction.

Due to the features of large actuation force, low-powerconsumption, and room temperature operation, our micropumpwithout the needs of complicated microfabrication and mechan-ical moving parts is suitable for the development of low-powerconsumption and compact micropumps for various applica-tions. One of the long-term goals of this paper is to develop alow-power-consumption micropump in a closed loop for microdirect methanol fuel cell (µDMFC) application. Meng et al.also recently proposed an electrolysis-bubble-actuated micro-pump for µDMFC application [20]. They took advantage ofthe bubbles traveling from hydrophilic to hydrophobic envi-ronment in a microchannel to move liquid to generate a netpumping flow. In that report, the pressure head was measuredat about 120–195 Pa, and the pumping rate was characterizedas 20–65 nl/s in the microchannel of 600 × 300 µm. Due tothe traveling feature of electrolytic bubbles, either choking orsticking phenomenon of electrolytic bubbles might appear inthe microchannel. In our micropump, the pumping mechanismdoes not rely on the long-distance traveling of electrolyticbubbles. Our proposed micropump could resolve disadvantagessuch as the complicated time-sequential power control, theneed of long nozzle-diffuser structure, and the choking/stickingphenomena of electrolytic bubbles in a microchannel. The the-oretical analysis, design, synthesis, microfabrication process,operating principles, and experimental results are describedas follows.

II. DEVICE OPERATION PRINCIPLE AND

MICROFABRICATION PROCESS

A. Operation Principle

Fig. 1(a) illustrates the design concept and operation princi-ple of our micropump, which consists of platinum electrodes, ahydrophilic microchannel, and the hydrophobic lateral breatherconnected to ambient air for the purpose of removing theelectrolytic bubbles. The hydrophobic lateral breather, whichis made of a top polydimethylsiloxane (PDMS) cover and theTeflon patterns extending to the outside of the microchannel onthe bottom substrate, not only degases the electrolysis bubblesbut also prevents the liquid from leaking out. Because of theroughness gradient design on the lateral breather, our microp-ump obtains a net pumping flow after the bubbles are degassedout of the microchannel. In our micropump design, we takeadvantage of the hydrophobic property variation of the liquid

Fig. 1. Illustration of the design concept and the pumping principle.(a) Schematic 3-D view of our proposed micropump. A net pumping flow alongthe x-direction of the microchannel in one pumping cycle via three processes[(b) bubble generation, (c) degassing, and (d) liquid movement] is illustratedfrom the side views of the micropump. The net pumping fluid is achieved overmultiple electrolysis bubble actuation cycles in the microchannel. Here, Pb,PL, and PR are the pressures. θr,L, θr,R, and θb are the contact angles.

on different rough hydrophobic surfaces to design a roughnessgradient surface that is located at the lateral breather. Moredetails about the roughness gradient design will be describedin Section III.

The pumping principle of our micropump relies on thesurface tension effect and the generation of multiple electrolysisbubble actuation cycles. The surface tension effect is controlledvia the periodic generation of electrolytic bubbles and theroughness gradient design of the microchannel surface, whichresults in the contact angle variation of liquid along the mi-crochannel. Fig. 1(b) and (d) illustrates the pumping principle

CHENG AND LIU: ELECTROLYSIS-BUBBLE-ACTUATED MICROPUMP BASED ON THE ROUGHNESS GRADIENT DESIGN 1097

of our micropump in one pumping cycle from the side viewof the micropump. The actuation mechanism in one pumpingcycle could be roughly divided into three processes: 1) bubblegeneration; 2) degassing; and 3) liquid movement. First, thebubble is generated by the voltage applied on the electrodes topush the liquid both forward and backward in the microchannel,as shown in Fig. 1(b). In the process of bubble generation,there is no net pumping flow. Next, the bubble vents out ofthe microchannel through the hydrophobic lateral breather tothe ambient air when the liquid between the two platinumelectrodes turns into electrolytic bubble completely, as shownin Fig. 1(c). Due to our roughness gradient design, the apparentcontact angle on the leading meniscus is larger than that onthe trailing meniscus, i.e., θr,R > θr,L > 90◦. Therefore, thepressure on the trailing meniscus of the liquid PL is largerthan the pressure on the leading meniscus of the liquid PR

according to the Laplace–Young’s equation [16]. Finally, bothliquid meniscuses flow back with different velocities to fill inthe empty section by the capillary forces during the degassingprocess, as shown in Fig. 1(d). As a result, one pumping cycleis completed, and a net pumping flow along the x-direction isachieved. The net pumping flow can be achieved over multipleelectrolysis bubble actuation cycles in the microchannel. Thenet volume displacement of liquid and the pumping rate aredominated by the geometry design of the microchannel, thedesign of the roughness gradient surface, the frequency and am-plitude of the applied voltage, and the design of the electrodes.All details will be described with the experimental results laterin this paper.

B. Microfabrication

The result of the microfabrication process for our micropumpis illustrated in Fig. 1(a). First, a standard p-type 〈100〉 siliconsubstrate is grown with the thermal oxide of 6000 Å that islithographically patterned and etched later as the oxide mask.Then, the substrate is etched by the wet etching process inKOH solution to define the microchannel. This silicon substrateis then grown with another 8000 Å thermal oxide to form thehydrophilic layer on top of the microchannel substrate. Then, a200 Å/1500 Å titanium/platinum electrode layer is evaporatedand patterned as the electrodes by a liftoff process. Next, TeflonAF (Amorphous Fluoropolymer, 1%, Dupont) is spin coated onthe wafer and cured at 180 ◦C for 6 h in a high-temperatureoven. The Teflon film is lithographically patterned and etchedas hydrophobic regions by the O2 plasma process to serve asthe bottom part of the hydrophobic lateral breather. The topPDMS cover with the roughness gradient structure is fabricatedby using a mold that is fabricated on another silicon waferwith the negative-tone UV photoresist (MicroChem Corpora-tion, SU-8 35). In this process, a curing agent and a PDMSprepolymer (Sylgard 184 Silicone Elastomer Kit) are mixed ina 1 : 10 weight ratio and stirred to ensure complete mixing. Theprepolymer mixture is degassed in a vacuum oven for removingair bubbles. Then, the prepolymer mixture is poured onto themold and cured at 65 ◦C for 1 h in the oven. After curing, thePDMS is peeled off from the mold. Finally, the top PDMS coveris bonded onto the top of the silicon wafer.

Fig. 2. Droplet on the surface forms: (a) the complete wetting and (b) thecomposite wetting based on the hydrophilic and the hydrophobic properties ofsubstrate surfaces, respectively.

III. THEORETICAL ANALYSIS AND MODELING

A. Roughness Gradient Design of Hydrophobic Surface

The contact angle has been commonly used to representsurface wettability. Surface wettability is a function of sur-face roughness. The latest experimental results confirm thatwettability can be tuned by surface roughness [25], [26]. Theearliest literatures reported that the contact angle of a dropleton a rough surface could be predicted by two main theoriesrelating the surface structure to the apparent contact angle. Thefirst theory was proposed by Wenzel [27], which assumes thatthe liquid completely wets the solid structure, as illustrated inFig. 2(a). The second theory was proposed by Cassie and Baxter[28], which assumes that the liquid does not wet the valleysof the structure and forms a composite surface on the roughsubstrate, as illustrated in Fig. 2(b). Afterwards, Bico et al. [26]fabricated the substrates with specific roughness and comparedthe measured contact angles with the prediction results. Theyclaimed good agreement and proposed that Wenzel’s formula isvalid for the hydrophilic surface and that Cassie and Baxter’sformula is valid for the hydrophobic surface.

Shastry et al., who first utilized the roughness gradient struc-ture to move the droplet, have also recently presented a modelfor the forces acting on the droplet resting on a roughness-controlled contact angle gradient [29]. A regular 2-D array ofsquare pillars creates the rough surface with the controlledroughness parameters φ and r, which are defined by the dimen-sion of square pillars (a× a), the pillar spacing (b), and thepillar height (c), i.e.,

φ =a2

(a+ b)2(1)

r =4bc+ (a+ b)2

(a+ b)2(2)

where φ is the surface area fraction, and r is the roughnessratio of the planar surface. The parameter φ determines theapparent contact angle on a composite surface based on Cassieand Baxter’s formula. r mainly correlates the relative stabilityof a droplet in the composite state. The stability of a droplet inthe composite state depends on the design of the parameters φand r, which will be discussed later.

The structural design of the surface roughness gradient inour device is illustrated in Fig. 3. The surface roughness isdominated by φ, which varies with the pillar patterns on thePDMS cover. For our roughness gradient design, φ decreases


Fig. 3. Illustration of the roughness gradient design on the hydrophobicsurface of the top PDMS cover that is made of ai × ai square pillars and pillarspacing bi arranged in a regular array.

along the x-direction of the microchannel, i.e., φ1 > φ2 >· · · > φn−1 > φn. The design of the roughness gradient surfacecould be modeled based on Shastry’s work. The roughnessgradient surface of our micropump is hydrophobic, which leadsto the formation of a composite surface. Thus, the roughness-controlled contact angle gradient could be predicted under theCassie–Baxter assumption

φi =a2

i

(ai + bi)2, i = 1, 2, 3, . . . , n (3)

cos θr,i = −1 + φi(cos θ + 1) (4)

where φi represents the surface area fraction of region i, whichis made of the rough surface on the PDMS side with ai × ai

square pillars and pillar spacing bi; θr,i is the apparent contactangle on the PDMS side in region i; and θ is the equilibriumcontact angle of the liquid droplet on the flat PDMS surface.The relationship between the apparent contact angle and thesurface area fraction could be theoretically derived via (4).Besides, previous report from Shirtcliffe et al. has also shownthat the dimension variation of square pillars allows the lengthof the contact perimeter per unit area to be varied withoutvarying the contact area per unit area [30]. As a result, thereis no change on the contact angles. Thus, the apparent contactangle is expected to remain constant when the surface areafraction is constant. Therefore, the surface area fraction is acritical factor for the roughness gradient design of hydrophobicsurface in our proposed micropumps. Besides, the previousliterature has also shown that a droplet on a rough surface mighttransit from composite state to complete state [29]. For thestability of the composite state, the following inequality holds:

cos θr <φ− 1r − φ

(5)

where θr is the apparent contact angle for the case of compositestate. This means the stability of a droplet in the composite stateis determined by the parametric design of φ and r, which arethe surface area fraction and the roughness ratio of the planarsurface, respectively. In this paper, the theoretical derivationsare based on the Cassie–Baxter assumption. Thus, the geometry

designs of pillars on the roughness gradient surface have tofollow (5). More details regarding the relationship between thepressure on the meniscus of liquid and the surface area fractionwill be derived based on the theoretical model addressed next.

B. Electrolysis

When an electric current is sent through the two noble metalelectrodes (such as platinum) in water, electrolysis takes place.The minimum equilibrium potential of hydrogen–oxygen elec-trolysis E0 is 1.23 V. In the electrolysis reaction, the oxygengas is produced at the anode, and the hydrogen gas is producedat the cathode, i.e.,

anode : 2H2O → 4H+ + 4e− + O2(g)

cathode : 2H2O + 2e− → 2OH− + H2(g). (6)

Under the assumption that all generated gases (O2 and H2)evolve in the form of gas bubbles, the total gas volume linearlydepends on the input electrical charge [17]. The total gasvolume generated by the electrolysis in the process of bubblenucleation could be estimated according to Faraday’s law ofelectrolysis and the ideal gas law [31]

N =It

zF(7)

PV =NRT (8)

where N is the moles of produced gas, I is the applied current,z is the number of excess electrons, F is Faraday’s constant(9.649 × 104 C/mol), t is the period of electrolysis, T istemperature, P is the ambient pressure, V is the volume of thebubble, and R is the gas constant (8.314 J K−1/mol). Under theassumption of constant temperature and atmospheric pressure,the volume of produced gases is proportional to the suppliedelectric current [32]. The theoretical bubble growth rate couldbe calculated by

dV

dt=

RIT

FPz. (9)

In our micropump, the bubble volume could not be preciselyestimated by Faraday’s law of electrolysis because this mi-cropump might have some loss of the bubble volume duringbubble expansion. However, the bubble growth rate is roughlyproportional with the applied current via raising the amplitudeof the applied voltage. The experimental results follow theproportional trend, which will be addressed in Section V.

C. Theoretical Modeling of Micropump

The electrochemical pump relies on the pressure increase viathe generation of oxygen and hydrogen gases when the appliedcurrent passes through the electrolytic liquid. The theoreticalmodels based on early literatures are addressed and summarizedas follows for the purpose of roughness gradient design in ourproposed micropump. The internal liquid pressures could beindividually modeled for the expansion period and the ventingperiod of the electrolytic bubble within one pumping cycle.


First, while the electrolytic bubble grows, the electrolytic bub-ble pressure Pb, as illustrated in Fig. 1(b), is modeled accordingto Faraday’s law of electrolysis and the ideal gas law as

Pb =NRT

V=

IRTt

zFV. (10)

While the electrolytic bubble is venting, the meniscuses of themoving liquid are driven based on the surface tension effect.A schematic illustration for the driving mechanism of ourproposed micropump is shown in Fig. 1(d). In the general case,one of the meniscuses of the liquid moves a forward displace-ment dx starting from the position x in the microchannel, onwhich the pressure could be modeled by using the total surfaceenergy approach [16], [33]. The effect of gravity is neglectedbecause of the shallow microchannel in our device. The surfaceenergies are related to the equilibrium contact angle governedby Laplace–Young’s equation

γsa = γsl + γla cos θr (11)

where θr is the apparent contact angle, and γsl, γsa, and γla

are the surface energies per unit area of solid–liquid, solid–air,and liquid–air, respectively. The differential interfacial energyof the system dUT could be derived as

dUT = γsldAsl + γsadAsa + γladAla

=(γsl − γsa)dAsl + γladAla

= γla(−cos θrdAsl + dAla) (12)

where Asl, Asa, and Ala are the interface areas of solid–liquid,solid–air, and liquid–air, respectively. The second equality in(12) is derived based on dAsa = −dAsl because the summationof Asl and Asa is a constant.

In our prototype device, micromachined PDMS pillars on topof the microchannel are employed to control the roughness ofthe hydrophobic surface. The bottom surface of the microchan-nel is coated with a hydrophilic layer of silicon. Thus, thedifferential interfacial energy on the trailing meniscus of theliquid could be modeled as

dUT = dUcover + dUmicrochannel

=[γla(−cos θr,i(PDMS)dAsl + dAla)

]PDMS

+[γla(−cos θr(SiO2)dAsl + dAla)

]SiO2

= γla [−wφi cos θPDMS + w(1 − φi)] dx

+ γla [−(2h+ w) cos θSiO2 ] dx

= γla [−wφi cos θPDMS + w(1 − φi)

− (2h+ w) cos θSiO2 ] dx (13)

where dUcover and dUmicrochannel are the differential interfacialenergies related to the top cover and the bottom microchannelsubstrate, respectively. θr,i(PDMS) is the apparent contact angleof liquid in region i on the PDMS side. θi(SiO2) is the apparentcontact angle of liquid on the thermal oxide film. θPDMS andθSiO2 are the equilibrium contact angles of the liquid drop onthe PDMS film and on the thermal oxide film, respectively.h and w are the depth and width of the microchannel, respec-

Fig. 4. Relationship between pressure and surface area fraction θ for differentmicrochannel depths (10, 20, and 50 µm). The microchannel width is fixed tobe 100 µm in all these calculations.

tively. To simplify the equation derivation, the microchannelwidth is assumed to be the average width of the nonorthogonalmicrochannel, which results from the wet etching process. Theshapes of meniscuses are also assumed to be constant curves.All the contact angles are assumed to be equilibrium contactangles neglecting the contact angle hysteresis. The pressure PL

on the trailing meniscus of the liquid moving in region i couldbe derived as [33]

PL = −dUT

dV̄

= −γla

wh[−wφi,L(1 + cos θPDMS)

+ w − (2h+ w) cos θSiO2 ] . (14)

The pressure PR on the leading meniscus of the liquid movingin region j could be obtained via a similar derivation

PR =− γla

wh[−wφj,R(1+cos θPDMS)+w−(2h+w) cos θSiO2 ]

(15)

where V is the liquid volume. PL and PR depend on the contactangles θPDMS and θSiO2, the geometry of the microchannel, andthe surface area fractions φi,L and φj,R. Here, φi,L and φj,R

are the surface area fractions in the roughness pattern region iand j, where the trailing and leading meniscuses of the liquidmove in, respectively. The pressure difference between thesetwo meniscuses, which affects the liquid displacement, couldbe derived as

∆P =PL − PR

= − γla

wh[−wφj,L(1+cos θPDMS)+wφj,R(1+cos θPDMS)]

=γla

h(φi,L − φj,R) [1 + cos θPDMS] . (16)

In our calculations, different microchannel depths (10, 20,and 50 µm) and the same microchannel width of 100 µmare assumed and used. Under the assumption of θSiO2 = 20◦,θPDMS = 110◦, and γla = 72 dyn cm−1, the relationship be-tween pressure and surface area fraction φ is derived and shownin Fig. 4 based on (14) and (15). The positive sign of thepressure means a forward pulling pressure. The calculationresults shown in Fig. 4 provide us the information of pressuredistribution versus different roughness gradient surfaces for


Fig. 5. Schematic diagram of the experimental setup. The outlet of our microchip is connected via a tube to the beaker, which could be raised by a high-precisionstage. The inlet is connected to a fixed beaker.

the design of our micropumps. In our design, the surface areafractions decrease along the x-direction of the microchannel,i.e., φi > φj for i < j. Therefore, the pressure difference isalways positive ∆P > 0 in our proposed micropump design.In other words, the micropump develops a net pressure headtoward the x-direction of the microchannel under the zero-net-flow condition. Thus, a net pumping flow is achieved via ourroughness gradient design.

The pressure difference in our micropump is dominated bythe roughness gradient design. φmax and φmin represent themaximum and the minimum of the surface area fraction, respec-tively, on the roughness gradient surface in the microchannel.The maximum pressure difference could be modeled as

∆Pmax(φmax, φmin, h) = PL(φmax, h) − PR(φmin, h). (17)

In this paper, micropumps with different roughness gradientdesigns and microchannel depths are implemented and char-acterized to demonstrate the performance. The experimentalresults will verify that the maximum pressure difference of themicropump dominates the liquid displacement in one pumpingcycle as well as the pumping flow rate. To simplify the deriva-tion, the model used in the above derivations does not includethe effect of contact angle hysteresis, which often results in asignificant portion of the surface tension effect for the dropletmotion [34], [35]. This effect depends on surface conditionof the microchannel. The calculated pressure will become lesswhen the contact angle hysteresis is taken into account.

D. Applied Voltage

The pumping flow rate relies on many factors, such as the ap-plied voltage, the duty cycle, and the driving frequency. These

imply that the actuation pulses play a dominant role for themaximum pumping flow rate. Besides, the expansion periodand the venting period of the bubble in one pumping cycle arealso critical parameters to regulate the square-wave actuationpulses. The operation frequency f and the duty ratio d aredefined as

f =1

texpand + tvent(18)

d =texpand

texpand + tvent(19)

where texpand and tvent are the expansion period and theventing period of the electrolytic bubble in one pumping cycle.The expansion period texpand is dominated by the period of theapplied voltage in one pumping cycle. The venting period tvent

is dominated by the bubble volume and the pressure magnitudeson both meniscuses.

IV. EXPERIMENTAL SETUP

To characterize the performance of our micropumps with dif-ferent parametric designs and operation conditions, the experi-mental setup is schematically represented in Fig. 5. To actuateour micropump, a Labview graphical program is implementedwith the programmable DAQ card (PCI-6534, NI) to regulatethe square-wave pulses and control the operation frequency andduty ratio. A direct-current power supply (E3631A, Agilent) isused to control the amplitude of the applied voltage throughthe electric control circuitry. A high-impedance field-electrictransistor in this electric control circuitry is used to performthe required voltage pulses. A relay is also used in this controlcircuitry to eliminate the alternating polarization and avoid thecorrosion of noble metal electrodes. An optical microscope


with a charge-coupled device (CCD) camera is utilized toobserve and record the pumping process of our micropumps.The open-loop micropump is fabricated for performance char-acterization. Two glass tubes are attached to the through-holesat the inlet and outlet of this micropump. The inlet and outletare connected to two beakers via soft tubes of 1.5-mm diameter.The beaker connected to the outlet could be accurately raisedby a high-precision stage. Pressure transducers (Lutron) with aresolution of 20 Pa are used to monitor the pressures.

The pumping pressure and the flow rate are characterizedvia the setup shown in Fig. 5. After the liquid of the beakerconnected to outlet tube is raised to the same height as theinlet meniscus and stabilizes, the driving voltage is appliedto start the pumping function. The meniscuses movement andthe bubble generation/degassing during the whole process arerecorded by a digital video system. For the flow rate char-acterization, two kinds of particles with different diametersof 3 ∼ 10 µm and 10 ∼ 30 µm (Polysciences, glass beads of2.48 g/cm3) are, respectively, mixed with water that are usedas the pumping liquid for two different micropumps with themicrochannel depths of 20 and 50 µm. The conductivity andthe pH value of pumping water are characterized as 0.35 ∼0.65 mS/cm and 6.5, respectively, by using the conductivity/pHmeter (EUTECH, PC510). By tracing the particles during atime period, the flow rate is calculated based on the averagespeed of the moving particles. The pumping flow rates of themicropumps are measured in open microchannels (i.e., negligi-ble back pressure) and characterized via applying the square-wave voltages of the specific voltage frequency and duty ratio.To approach the maximum flow rates, the expansion periodsand the venting periods are also characterized via the recordedimages under different applied voltages.

For the characterization of the pressure head, the beakerconnected to the outlet is lifted to adjust the pressure head. Inthese measurements, the pumping rates are recorded with thechange of pressure head, which is monitored via the pressuremeter. The maximum pressure head is recorded when theliquid stops flowing forward and starts flowing backward. Inaddition, the leakage pressure is characterized by sealing oneof two inlet/outlet ports. The leakage pressure is recorded viaobserving the phenomenon that the liquid wets the breathinghole of the lateral breather during the lifting process of thebeaker connected to the outlet.

For contact angle measurements, a small-volume drop ofwater (∼5 µL) is gently placed on the experimental surface.A photograph of the side profile of the liquid drop is taken byusing a CCD camera with adjustable optical focus lens. Theradius of the spherical cap is measured by processing the side-view image to obtain the free surface of the spherical cap. Then,we fit it with a circle to get the radius of the spherical droplet.The equilibrium contact angle θ is calculated from the radius ofthe fitting circle R and the height of the droplet H by using theformula θ = cos−1((R−H)/R).

V. EXPERIMENTAL RESULTS AND DISCUSSION

Fig. 6(a) shows a prototype device that has two micropumpswith different roughness gradient design on a chip, which is

Fig. 6. Experimental demonstration for the pumping function of our micro-pump. (a) Prototype device with two micropumps on a chip, which is madeof the PDMS cover (15 mm × 11 mm × 3 mm) and the silicon substrate(15 mm × 15 mm × 0.5 mm). The time-sequential pictures captured from thetop view of the micropump show the processes of pumping flow in one pumpingcycle including (b) bubble generation, (c) degassing, (d) liquid movement, and(e) net pumping flow after the electrolytic bubble is degassed completely. Thecross section of microchannel is 100 µm × 20 µm.

made of the PDMS cover (15 mm × 11 mm × 3 mm) andthe silicon substrate (15 mm × 15 mm × 0.5 mm). Table Isummarizes the distribution of surface area fraction (φi) forour prototype micropumps with four different structure designs.The microchannel width w is fixed to be 100 µm for all thesemicropumps. The micropumps are operated under the conditionof open microchannels for pumping characterization. The time-sequential pictures, which are captured from the top view ofthe micropump, in one pumping cycle for micropump#1 arerecorded as shown in Fig. 6(b) and (e). After the microchan-nel is filled with the sample liquid, the electrolytic bubbleis generated in the active region by applying an alternatingelectric current to the platinum electrodes. The liquid is then


TABLE IDESIGN DETAILS FOR FOUR TYPES OF MICROPUMPS WITH DIFFERENT ROUGHNESS GRADIENT DESIGNS (ΦA AND ΦB).

THE MICROCHANNEL WIDTH (w) IS FIXED TO BE 100 µm FOR ALL THESE MICROPUMPS

Fig. 7. Experimental characterization for the apparent contact angle θr on thehydrophobic surface versus different surface area fraction φ. The dots (�) andthe solid line ( ) represent the experimental results and theoretical predictionbased on Cassie–Baxter’s theory, respectively.

separated into two sections and pushed toward both inlet andoutlet directions, respectively, in the microchannel, as shownin Fig. 6(b).

Fig. 6(c) and (d) shows the processes when the appliedvoltage on the platinum electrodes is turned off. The electrolysisbubble is degassed through the lateral breather to ambient air.Both meniscuses of liquid then flow back with different veloci-ties to fill in the empty section of liquid in the microchannel. Asa result, a net pumping flow with the function of degassing theelectrolytic bubbles is achieved in one pumping cycle based onthe roughness gradient design, as shown in Fig. 6(e). Therefore,liquid pumping is achieved over multiple electrolysis bubbleactuation cycles in the microchannel.

A. Characterization of Contact Angles

To characterize the contact angles for various surface areafractions (φ), several different testing microfluidic chips arefabricated. All of these microfluidic chips have a design sim-ilar to our proposed electrolysis-bubble-actuated micropumpexcept for no roughness gradient on the PDMS covers. EachPDMS cover (25 mm × 25 mm × 3 mm) has pillars with thesame surface area fraction (φ) along the microchannel. Allof these testing chips have different surface area fractionsvarying from 0.1 to 0.9 via the position arrangement of pillars.Fig. 7 shows the characterized and the theoretical relationshipbetween the apparent contact angle θr and the surface areafraction φ. The experimental characterization agrees with thetheoretical prediction from Cassie and Baxter’s theory. Here,the characterized data are taken from the average results of five-

Fig. 8. Characterization of leakage pressure for five hydrophobic lateralbreathers with different spacing between adjacent pillars (5, 10, 20, 30, and40 µm).

times experiments because it is difficult to precisely measurethe equilibrium angles under the existence of contact anglehysteresis on these microchannel surfaces.

B. Leakage Testing

The capability of the micropump to be operated againstback pressures is one of the key requirements for operating amicropump in a closed-loop environment. In our micropump,the liquid is pumped forward based on the breathing function ofthe hydrophobic lateral breather in our micropumps. When theback pressure is higher than the leakage pressure of the lateralbreather, the breathing of gases would be blocked to make ourmicropump lose the ability of degassing electrolytic bubbles.The leakage pressure, which represents the maximum allowedliquid pressure without leakage through the breathing holes, isdominated by the dimension of breathing hole at the lateralbreather and follows Laplace–Young’s equation [36]. In ourcase, the leakage pressure of breather is affected by the spacingbetween pillars. Thus, it is important to define the criticalspacing that could support a given liquid pressure withoutblocking the breather. For the leakage pressure measurements,five rough surfaces with the same pillar dimension (25 µm ×25 µm × 30 µm) and different spacing between adjacent pillars(5, 10, 20, 30, and 40 µm) are fabricated to form different lateralbreathers. The experimental results for the leakage pressureversus the spacing between adjacent pillars are shown in Fig. 8.The solid line is obtained by a linear fitting of the measureddata. The maximum leakage pressure of 17.9 kPa is achievedin our prototype micropumps based on the pillar spacingof 5 µm.


Fig. 9. Net pumping displacement of liquid in one pumping cycle versus thecalculated maximum pressure difference for the four types of micropumps. Thedesigns and specifications for the four types of micropumps with different mi-crochannel depths and surface roughness gradients are summarized in Table I.

C. Pressure and Flow Rate Testing

According to (14)–(17) and the theoretical derivation inFig. 4, the maximum pressure difference is dominated by theroughness gradient design and the microchannel depth. Fourtypes of micropumps with the designs of different roughnessgradients (ΦA and ΦB) and different microchannel depths(20 and 50 µm) as summarized in Table I are used to evaluatethe performance of our proposed micropumps. Fig. 9 shows therelationship between the calculated maximum pressure differ-ence and the net pumping displacement of liquid in one pump-ing cycle for these four types of micropumps. Compared withother prototype micropumps, micropump#2 with the largestmaximum pressure difference pumps the liquid to achieve thelargest displacement in one pumping cycle. The large maximumpressure difference is achieved by the design of the shallowmicrochannel depth and the large surface roughness gradient.Here, the pumping displacement of the liquid is characterizedvia measuring the average displacement of the moving particlesunder the initial condition of zero flow. The data for the pressuredifference shown in Fig. 9 are derived theoretically based on(14)–(17) and Fig. 4. The error bars shown in Fig. 9 result fromthe particle displacement variation among ten measurements foreach pressure difference design.

For the characterization of the maximum volume flow rate,micropump#2 and micropump#4 with the same design of largesurface roughness gradient (ΦB) and different microchanneldepths (20 and 50 µm) are used to observe the effect ofmicrochannel depths. To approach the optimal pumping ratesunder different applied voltages, the expansion period and theventing period in one pumping cycle are characterized first.Fig. 10 shows the expansion period of the electrolysis bubbleto fill the roughness gradient region in the microchannel underdifferent applied voltages (7, 9, 11, 13, and 15 V). The curvesare obtained by simple linear fitting among the measured data.The characterized venting periods are approximately 100 and150 ms for micropump#2 and micropump#4, respectively.

Based on the characterization results of the expansion periodtexpand and the venting period tvent, the operation frequency fand the duty ratio d of the applied voltage could be optimized toenhance the flow rate of micropumps. Fig. 11 shows the charac-terization results for these enhanced flow rates of micropump#2

Fig. 10. Expansion period of the electrolysis bubble texpand to fill theroughness gradient region in the microchannel versus the applied voltage formicropump#2 and micropump#4.

Fig. 11. Experimental results for the enhanced pumping volume flow rateversus the operation frequency under different amplitudes of applied voltages(7, 9, 11, 13, and 15 V).

and micropump#4. The data are approximated via the curvefitting of two second-order polynomial functions. The error barsshown in Fig. 11 result from the flow rate variation among tenmeasurements. The applied voltage signals are regulated bythe operation frequency and duty ratio, which are calculatedvia (18) and (19). The optimal frequency of applied voltagewith respect to the pumping flow rate is dominated by theexpansion period and the venting period of electrolysis bubble.With the applied voltage of 15 V at 4.5 Hz, a maximumpumping rate of 114 nl/min is achieved for our micropump#2that has a microchannel cross section of 100 µm × 20 µm.These results indicate that the flow rate in our micropumpscould be simply regulated by adjusting the voltage pulses toapproach the appropriate flow rate in the microfluidic chip.

The characterized relationship between flow rate and pres-sure head is shown in Fig. 12. The flow rate is observed in thepresence of different elevations between the inlet and the outletto generate a pressure difference. The characterized data shownin Fig. 12 for micropump#2 and micropump#4 are measuredunder the applied voltages of 15 V at 4.5 and 2.5 Hz, respec-tively. When the flow rates decay to zero, the pressure headsare characterized as 1.9 and 0.8 kPa for micropump#2 andmicropump#4, respectively. When the pressure head is closeto zero, the maximum flow rates, which are 114 and 95 nl/minfor micropump#2 and micropump#4, respectively, are achieved.Besides, the power consumption Pw and the thermodynamic ef-ficiency η are also important parameters. The maximum power


Fig. 12. Characterized relationship between the pumping volume flow rateand the pressure head for micropump#2 and micropump#4. The details for thestructure designs of micropump#2 and micropump#4 are listed in Table I.

consumption of the micropump is measured to be 1.25 mW foran applied voltage of 15 V, which has a corresponding averagecurrent of 83 µA during the pumping operation. Because ofthe approximately linear relationship between the flow rate andthe pressure head, the thermodynamic efficiency η could beestimated based on the power consumption, the maximum flowrate Qmax, and the maximum pressure head (Phead)max byusing the formula η = (1/4)(Qmax(Phead)max/Pw) [1]. Whenour micropump (micropump#2) is operated at f = 4.5 Hzand Pw = 1.25 mW, this micropump produces a Qmax of114 nl/min and a (Phead)max of 1.9 kPa, as shown in Fig. 12.The thermodynamic efficiency is about η = 7.4 × 10−6%.

D. Corrosion

The direct application of voltage pulse trains forces the elec-trodes to continuously alternate polarization during the opera-tion. Alternating-current polarization has been reported to leadto the corrosion and the roughening of noble metal electrodes[37]. Specifically, noble metals such as platinum and palladium,which readily dissolve hydrogen, exhibit rapid corrosion rates.Howe et al. reported that the corrosion of noble metal electrodescould be avoided by alternating the driving voltage betweena fixed operation potential and an unbiased condition duringoperation [21]. To enhance the micropump performance andimprove the lifetime operation, the driving voltage is alternatedbetween a fixed operating potential and an unbiased voltageground during the operation via a relay to minimize the metalcorrosion. Although the long-term test has not been performed,we have applied the current to the electrodes continuously formore than 3 h without any observable electrical degradation.

VI. CONCLUSION

A novel electrolysis-bubble-actuated micropump with thedesign of the roughness gradient on the microchannel hy-drophobic surface and the lateral breather has been successfullydemonstrated and reported in this paper. Theoretical modelsare addressed in this paper to study the pumping principle,which is dominated by the pressure difference between two

meniscuses of the liquid in our proposed micropump. Thedesign, synthesis, micromachining process, and comprehensivecharacterization of our micropumps have also been presented.Furthermore, experimental results successfully demonstrate thepumping function of our micropumps with different designsof the roughness gradient and the microchannel depth. Thepressure head of 1.9 kPa and the pump rate of 114 nl/minare measured for our micropump with the microchannel crosssection of 100 µm × 20 µm. Compared to other reported mi-cropumps for µDMFC application, our low-power consumptionmicropump has the potential to be applied in a closed loopfor µDMFC application. Further work to modify, optimize,and integrate this electrolysis-bubble-actuated micropump forapplications like µDMFC and lab chips is under going inour group.

The features of our proposed micropumps on compact size,simple microfabrication, low-power consumption, and roomtemperature operation make it promising to be integrated withother multiple components to form microfluidic systems forapplications such as lab on a chip, biochip, drug delivery,and µDMFC. Because of the feature of room temperatureoperation, this micropump is specifically suitable for variousbio-applications. However, the performance of this micropumpcritically depends on the PDMS surface maintaining its intrinsiccontact angle. In some specific cases, this micropump mightlose the pumping feature. For example, proteins would bindnonspecifically to the PDMS surface, which would change thecontact angle and even make it hydrophilic.

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Chih-Ming Cheng received the B.S. degree in me-chanical engineering from National Taiwan Uni-versity of Science and Technology, Taipei, Taiwan,R.O.C., in 1998, and the M.S. degree in mechanicalengineering from Taiwan University, Taipei, in 2000.He is currently working toward the Ph.D. degree atNational Tsing Hua University, Hsinchu, Taiwan.

He is currently with the Micro-Systems and Con-trol Laboratory, National Tsing Hua University,where he focuses on the development of microfluidicswitches for bio-analytical devices and micropumps

for micro direct methanol fuel cell (DMFC) applications.

Cheng-Hsien Liu (M’02) received the B.S. de-gree in power mechanical engineering from NationalTsing Hua University, Hsinchu, Taiwan, R.O.C., in1987, the M.S. degree in mechanical engineeringfrom Lehigh University, Bethlehem, PA, in 1992,and the M.S. degree in electrical engineering andthe Ph.D. degree in mechanical engineering fromStanford University, Stanford, CA, in 1995 and 2000,respectively.

While at Stanford, he worked with Dr. Kenny atthe Stanford Microstructures and Sensors Laboratory

and focused his Ph.D. work on high-performance tunneling MEMS sensors.In 1999–2000, he was a Senior Electrical Engineer at Halo Data DevicesInc., San Jose, CA, where he focused on the development of microdrivesfor portable information storage applications. Since Autumn 2000, he hasbeen with National Tsing Hua University, where he is currently an AssociateProfessor in the Power Mechanical Engineering Department. He also serves asthe Department Vice Chair and the Division Head of the Control Division. Hecurrently oversees graduate students in the Micro-Systems and Control Labora-tory, whose research activities cover a variety of areas such as biomimetic arraychip for bio-object manipulation targeting for tissue engineering/drug screeningapplications, liver Labchip, advanced tunable MEMS grating, and microsystemrobust control.

an electrolysis bubble actuated micropump

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