manufacturing of dielectric barrier discharge plasma actuator by nicole m. houser · 2013. 11....

Manufacturing of Dielectric Barrier Discharge Plasma Actuatorfor Degradation Resistance

by

Nicole M. Houser

A thesis submitted in conformity with the requirements

for the degree of Masters of Applied Science

Graduate Department of Aerospace Studies

University of Toronto

c© Copyright 2013 by Nicole M. Houser

Abstract

Manufacturing of Dielectric Barrier Discharge Plasma Actuator for Degradation

Resistance

Nicole M. Houser

Masters of Applied Science

Graduate Department of Aerospace Studies

University of Toronto

2013

The performance and broader application of dielectric barrier discharge (DBD) plasma

actuators are restricted by the manufacturing methods currently employed. In the current

work, two methodologies are proposed to build robust plasma actuators for active �ow

control; a protective silicone oil (PDMS) treatment for hand-cut and laid tape-based

actuators and a microfabrication technique for glass-based devices. The microfabrication

process, through which thin �lm electrodes are precisely deposited onto plasma-resistant

glass substrates, is presented in detail. The resulting glass-based devices are characterized

with respect to electrical properties and output for various operating conditions. The

longevity of microfabricated devices is compared against silicone-treated and untreated

hand-made devices of comparable geometries over 60 hours of continuous operation. Both

tungsten and copper electrodes are considered for microfabricated devices. Human health

e�ects are also considered in an electromagnetic �eld study of the area surrounding a live

plasma actuator for various operating conditions.

ii

Acknowledgements

First and foremost, I would like to thank Dr. Philippe Lavoie for granting me the

opportunity to complete my graduate studies at the University of Toronto Institute for

Aerospace Studies (UTIAS) in the Flow Control and Experimental Turbulence (FCET)

lab. I would also like to thank Dr. Craig Steeves for refereeing this thesis.

I would like to extend the deepest thanks to my friends in the FCET group and to

the UTIAS community for their cherished friendship and continuous support throughout

my studies. In particular, I owe a debt of gratitude to Dr. Ronnie Hanson who served

as a knowledgeable mentor and supportive role model for many in the FCET lab, espe-

cially myself. Much of this thesis builds from groundwork laid by Ronnie and truly, his

expertise was always appreciated. For sharing their plasma actuator wisdom with me, I

gratefully acknowledge Mr. Luke Osmokrovic, Dr. Arash Naghib-Lahouti, and Dr. John

Murphy. I would also like to thank Todd Simpson and especially, Tim Goldhawk of the

Western Nanofabrication Facility for their patience and immense help while making (and

sometimes breaking) the microfabricated actuators featured in the current work, as well

as their contributions to the SEM/EDS analysis in this thesis.

On a personal note, I wish to thank Ra�k for his treasured partnership. I also

appreciate the chance to thank my close childhood friends, Helen and Sonia; my favourite

physics pals, Shannon, Corey, and Julian; my sister, Katrina (& family); and my mom,

Deborah for being exactly as awesome as they are.

iii

Contents

1 | Introduction

1.1 Motivation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1

1.2 Organization of Thesis . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2

2 | DBD Plasma Actuator

2.1 DBD Plasma Actuator Physics . . . . . . . . . . . . . . . . . . . . . . . 3

2.2 Operational Trends . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5

2.3 DBD Actuator Optimization . . . . . . . . . . . . . . . . . . . . . . . . . 6

2.4 Applications to Flow Control . . . . . . . . . . . . . . . . . . . . . . . . 8

3 | DBD Plasma Actuator Manufacturing

3.1 Conventional Methods & Materials . . . . . . . . . . . . . . . . . . . . . 9

3.2 Device Degradation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10

3.3 Methods for Increased Device Longevity . . . . . . . . . . . . . . . . . . 11

3.3.1 Protective Coating . . . . . . . . . . . . . . . . . . . . . . . . . . 12

3.3.2 Microfabrication . . . . . . . . . . . . . . . . . . . . . . . . . . . 12

4 | Experimental Methods

4.1 CCD Camera Experimental Set-up . . . . . . . . . . . . . . . . . . . . . 17

4.2 Electrical Characterization . . . . . . . . . . . . . . . . . . . . . . . . . . 18

4.2.1 Power Consumption & Capacitance Calculations . . . . . . . . . . 18

4.2.2 Probe Capacitor Independence Check . . . . . . . . . . . . . . . . 20

4.3 Particle Image Velocimetry . . . . . . . . . . . . . . . . . . . . . . . . . . 21

4.3.1 PIV Experimental Set-up . . . . . . . . . . . . . . . . . . . . . . 21

4.3.2 Data Processing . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22

4.4 Electromagnetic Field Measurement Set-up . . . . . . . . . . . . . . . . . 22

iv

5 | Characterization of Microfabricated Devices

5.1 Actuator Speci�cations . . . . . . . . . . . . . . . . . . . . . . . . . . . . 24

5.2 Electrical Quanti�cation . . . . . . . . . . . . . . . . . . . . . . . . . . . 25

5.3 Momentum Transfer to Air . . . . . . . . . . . . . . . . . . . . . . . . . . 30

6 | Device Degradation Studies

6.1 Actuator Speci�cations . . . . . . . . . . . . . . . . . . . . . . . . . . . . 34

6.2 PMDS Treated Kapton Actuators . . . . . . . . . . . . . . . . . . . . . 34

6.3 Microfabricated Glass Actuators . . . . . . . . . . . . . . . . . . . . . . 40

7 | Electromagnetic Radiation from Plasma Actuators

7.1 Electric Field . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 46

7.2 Magnetic Field . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 48

7.3 Concerns with EMF Radiation Exposure . . . . . . . . . . . . . . . . . . 50

8 | Summary & Conclusions

8.1 Investigation of Microfabricated Devices . . . . . . . . . . . . . . . . . . 53

8.2 Degradation of Plasma Actuators . . . . . . . . . . . . . . . . . . . . . . 54

8.3 Electromagnetic Field Considerations . . . . . . . . . . . . . . . . . . . . 55

8.4 Concluding Remarks . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 55

References

v

List of Tables

2.1 Induced velocity and applied voltage relation for plasma actuators. . . . . 6

3.1 Spin-coat recipe speci�cations. . . . . . . . . . . . . . . . . . . . . . . . . 14

5.1 Summary of plasma actuators used in characterization of microfabricated

devices. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 25

6.1 Summary of actuators used in CCD degradation studies. . . . . . . . . . 34

7.1 Field strength limits set by Health Canada's Safety Code 6. . . . . . . . 51

vi

List of Figures

2.1 Schematic of DBD plasma actuator. . . . . . . . . . . . . . . . . . . . . . 3

2.2 Ion drift during and typical current response AC cycle. . . . . . . . . . . 4

2.3 Power consumption as a function of applied voltage for various studies. . 5

3.1 Photolithographic process for plasma actuators, 2D cross-sectional view. . 13

3.2 The top-view of the undercut created with a bilayer resist process after

metal deposition as seen through a microscope. . . . . . . . . . . . . . . 13

3.3 Various stages in microfabrication procedure: (a) spin-coat application of

LOR 30B, (b) the mask alignment tool used for UV exposure, (c) substrate

prepared for second side metal deposition, (d) four substrates loaded in

the sputtering chamber following metal deposition, (e) sputtering vapour

deposition interface, (f) metal lifting-o� from glass substrate in solvent to

reveal plasma actuator array, and (g) a microfabricated plasma actuator

device. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15

3.4 A 1mm interelectrode gap of a glass actuator exhibiting (a) partial dis-

charges in Kapton encapsulation layers during actuation and (b) the re-

sulting damage in insulating material. . . . . . . . . . . . . . . . . . . . . 16

4.1 Schematic of experimental set-up for degradation studies with CCD camera. 17

4.2 A typical Q− V cyclogram for determination of electrical quantities. . . 19

4.3 Probe capacitor independence check for a microfabricated copper on glass

actuator operated at 6 kV, 4 kHz. . . . . . . . . . . . . . . . . . . . . . . 20

4.4 (a)Schematic of experimental set-up and (b) actual set-up for PIV exper-

iments with laser �ring. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21

4.5 Schematic of experimental set-up for electric and magnetic �eld studies. . 22

5.1 Plasma actuator dimension de�nitions. . . . . . . . . . . . . . . . . . . . 24

5.2 Power consumption per unit length as a function of (a) frequency and (b)

applied voltage for microfabricated actuators (at 4 kHz). . . . . . . . . . 26

vii

5.3 Cold and e�ective capacitances per unit length as a function of (a) fre-

quency and (b) applied voltage (at 4 kHz) for microfabricated actuators. . 27

5.4 Charge across the electrodes in response to applied voltage for (a) a treated

Kapton actuator, (b) a copper on glass actuator, and (c) a tungsten on

glass actuator with (d) the mean cycle after 60 hours of continuous oper-

ation at 6 kV, 4 kHz. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 28

5.5 Variations in a Q− V cyclogram for microfabricated and handmade actu-

ators after 60 hours of operation at 6 kV, 4 kHz. The substantial di�erence

in cyclogram area between handmade and microfabricated actuators is due

to di�erences in actuator length. . . . . . . . . . . . . . . . . . . . . . . . 29

5.6 The e�ective capacitance values calculated from the positive half cycle

slopes of cyclograms (C+eff ), the negative half cycle slopes of cyclograms

(C−eff ), and using the histogram method for a microfabricated tungsten on

glass actuator over a period from 10 to 20 hours into continuous operation

at 6 kV, 4 kHz. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 29

5.7 Maximum induced velocity as a function of (a) power consumption (with

increasing voltages at 4 kHz), (b) frequency (at 6 kV), and (c) applied

voltage (at 4 kHz)for microfabricated actuators. . . . . . . . . . . . . . . 31

5.8 Velocity pro�les, normalized according to Equation 5.3, of �ow induced by

a actuators of various construction methods and materials,each operated

at 7.5 kV, 4 kHz. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 32

6.1 Variations in the (a) power consumption, (b) e�ective capacitance, and

(c) cold capacitance of various Kapton-based actuators during 60 hours

of continuous operation at 6 kV, 4 kHz. Values shown are normalized by

respective initial measurements at t=0hrs. . . . . . . . . . . . . . . . . . 35

6.2 CCD images of exposed electrode (top) and dielectric surface (bottom)

for (a)hand-laid copper on 0.18mm Kapton (from [18]), (b)hand-laid cop-

per on PDMS treated 0.18mm Kapton, and (c)hand-laid copper on twice

PDMS treated 0.18mm Kapton actuators during continuous operation at

6 kV, 4 kHz. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 37

6.3 Surface of PDMS treated Kapton-based actuator (0.18mm thick) following

60 hours of continuous operation at 6 kV, 4 kHz. . . . . . . . . . . . . . . 38

viii

6.4 SEM images of (a) degradation patterns in the dielectric surface of silicone

treated 0.18mm Kapton actuator and (b) initiation holes in the dielectric

surface of a 0.18 mm Kapton actuator with two oil treatments. Both actu-

ators were operated for 60 hours at 6 kV, 4 kHz. . . . . . . . . . . . . . . 39

6.5 Variations in the (a) power consumption, (b) e�ective capacitance (posi-

tive half cycle),(c) e�ective capacitance (negative half cycle), and (d) cold

capacitance of various actuators during 60 hours of continuous operation

at 6 kV, 4 kHz. Values shown are normalized by respective initial measure-

ments at t=0hrs. No data was recorded between t=32hrs and t=45hrs

for the microfabricated actuator with tungsten electrodes. . . . . . . . . . 41

6.6 CCD images of exposed electrode and dielectric surface for (a)microfabricated

copper on 0.3mm glass, and (b) microfabricated tungsten on 0.3mm glass

actuators during continuous operation at 6 kV, 4 kHz. . . . . . . . . . . . 42

6.7 Comparison of unused (left) and used (right) plasma actuators via SEM

magni�cation for (a) hand-laid Kapton and copper tape, (b) sputter de-

posited copper electrodes on glass, and (c) sputter deposited tungsten

electrodes on glass. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 43

6.8 Comparison of the interface between the plasma-forming electrode edge

(top) and the dielectric surface (bottom) for used microfabricated actua-

tors with (a) sputter deposited copper electrodes and (b) sputter deposited

tungsten electrodes via SEM magni�cation. . . . . . . . . . . . . . . . . . 44

6.9 Plasma generation response to increasing applied voltage by a used micro-

fabricated actuator with (a) sputter deposited copper electrodes and (b)

sputter deposited tungsten electrodes. The applied voltage amplitude is

listed above each image column in kV. . . . . . . . . . . . . . . . . . . . 45

7.1 The resultant electric �eld strength at all orientations in response to vari-

ous operating voltages (at 4 kHz). . . . . . . . . . . . . . . . . . . . . . . 47

7.2 The deviation of each measurement from the position averaged electric

�eld strength for all operating conditions. . . . . . . . . . . . . . . . . . . 47

7.3 The resultant electric �eld strength at 90◦ orientation as a function of

(a)radial distance from actuator centre for various operating conditions

and (b)operating voltage at various radial locations (at 4 kHz). . . . . . . 48

7.4 The resultant magnetic �eld strength at all orientations in response to

various operating voltages (at 4 kHz). . . . . . . . . . . . . . . . . . . . . 49

ix

7.5 The resultant magnetic �eld strength at 90◦ orientation as a function of

(a)radial distance from actuator centre for various operating conditions

and (b)operating voltage at various radial locations (at 4 kHz). . . . . . . 49

x

1 | Introduction

1.1 |Motivation

As increasing attention is paid to both the environmental and economical impacts of

aircraft operation, e�ective actuators that enhance vehicle performance have become an

important research focus. Dielectric barrier discharge (DBD) plasma actuators, herein

referred to as plasma actuators, have piqued the interest of the �ow control community

as a novel type of �ow control device for several reasons. These light-weight devices have

a low pro�le, can be laminated to aerodynamic surfaces, and provide fast response for

feedback control. Studies have demonstrated the applicability of plasma actuators to a

variety of �ow control situations such as transition delay [16, 20] and noise mitigation [26].

However, this technology is still in its infancy and conventional construction methods and

materials impose limitations on device performance due to signi�cant imperfections and

material wear. The need for improved actuator robustness is the primary motivation

behind this project which aims to improve manufacturing techniques and identify more

suitable materials for plasma actuators.

Health monitoring of actuator performance for robust systems-level integration was

identi�ed by Cattafesta and Sheplak [3] as a future direction that will facilitate the transi-

tion of laboratory-scale devices to full-scale applications. For successful implementation,

the limitations of �ow control actuators must be quanti�ed and addressed. In the present

work, limitations of conventional construction methods were addressed by the exploration

of a protective surface treatment for hand-made devices as well as the development of a

rigorous microfabrication procedure for plasma actuators with an emphasis on plasma re-

sistant materials. The electrical properties of these devices were evaluated over extended

actuation periods and compared against the most prevalent construction method found

in the literature. The objective of this work is to produce and characterize robust and

repeatable devices for experimentation as a stepping-stone to industrial applications.

1

Chapter 1. Introduction 2

1.2 |Organization of Thesis

Chapter 2 o�ers a description of DBD plasma actuators, followed by a review of oper-

ational trends from the literature, and an overview of various plasma actuator applica-

tions. Chapter 3 provides a detailed review of materials and methods used for plasma

actuator construction in previous studies, highlighting the major limitations encountered

with these conventional construction techniques. The means used to increase actuator

longevity developed for the present work, protective surface treatment and microfabrica-

tion, are also described in detail.

The results presented in this thesis were obtained using a variety of di�erent exper-

imental set-ups an procedures, all of which are outlined in Chapter 4. Chapter 5 serves

to characterize microfabricated devices with respect to electrical quantities and momen-

tum transfer performance. The degradation resistance of actuators fabricated with the

various methods and materials are compared over prolonged use in Chapter 6.

The electromagnetic �elds surrounding a plasma actuator are characterized in Chap-

ter 7 and the concerns to both researchers and laboratory equipment are stated.

Finally, Chapter 8 summarizes the main insights of this work and concludes with

suggestions for future e�orts to improve plasma actuator robustness and repeatability

for applications to �ow control.


2.1 |DBD Plasma Actuator Physics

grounded electrode

dielectric substrateplasmaexposed electrode

encapsulating substrate

didididididididielelectricplplasmatrode

Vac

induced flow

Figure 2.1: Schematic of DBD plasma actu-ator.

In their simplest con�guration, DBD

plasma actuators consist of two asymmet-

rically arranged electrodes separated by a

dielectric material, as shown in Figure 2.1.

The grounded electrode is encapsulated to

prevent the unwanted formation of plasma

on the underside of the actuator. The ex-

posed electrode is placed atop the dielec-

tric material. When a su�cient AC volt-

age is supplied, the asymmetric electrode

con�guration creates an electric �eld on

top of the dielectric that weakly ionizes the

air above the buried electrode forming a plasma discharge. The charged plasma experi-

ences a Lorentz force in the presence of the electric �eld causing a net body force in the

direction of decreasing electric �eld potential. This body force draws the ambient (neu-

trally charged) air towards the wall and expels �uid away from the exposed electrode.

The pressure drop above the electrodes from the ejection of �uid leads to a suction e�ect

pulling air towards the actuator. The ionized air generates low intensity violet-coloured

light emission.

Plasma actuators operate via continuous successions of electron avalanche mecha-

nisms in opposite directions over the two halves of the applied AC cycle. When the

exposed electrode is more negative than the dielectric surface, the avalanche growth of

electrons produced by secondary emission results in the deposition of negative charge on

the surface of the dielectric [12, 36]. The electrons are returned to the electrode on the

subsequent positive going half cycle of the discharge [46, 51]. The accumulation of charge

on the dielectric surface opposes the applied voltage as the ions in the plasma arrange

3

Chapter 2. DBD Plasma Actuator 4

to cancel the electric �eld. The discharge self-terminates on the surface of the dielectric.

Alternating current is required for continuous discharge, otherwise, the discharges will

choke as the electric �eld is quenched [9, 40, 46]. Figure 2.2, depicts the ion drift during

the AC cycle. During the negative-going half of the AC cycle, air ionizes at 2.2a until

voltage on exposed electrode stops becoming more negative at 2.2b. During the positive-

going half cycle electrons are emitted from the dielectric surface to exposed electrode at

2.2c. Air ionizes until voltage stops becoming more positive at 2.2d.

Vac

a

b

c

d

0.2 0.3 0.40.1

20

10

0

-10

-20

5

2.5

0

-2.5

-50.50

time [ms]

curren

t [mA

]

Vac

voltag

e [k

V]

Figure 2.2: Ion drift during and typical current re-sponse AC cycle.

It has been well documented

that plasma actuators have an

asymmetric current response to

symmetric AC signal, as shown

in Figure 2.2. For the negative-

going phase of the applied sig-

nal, discharges occur in rapid

succession in a di�use discharge

whereas the positive-going phase

generates highly �lamentary dis-

charges, fewer in number but with

greater intensity [8, 40]. This

di�erence in discharge behaviour

has led to ongoing debates con-

cerning the nature of momentum

transfer with respect to the dif-

ferent half cycles. Disagreement

exists about whether the plasma

strongly pushes (accelerates) the

�ow downstream during one half

cycle and weakly pulls (deceler-

ates) the �ow upstream in the op-

posite half cycle, or if the plasma pushes the �ow downstream during both half cycles.

The two main con�icting ideas are named the push-pull theory and the push-push

theory. Support for the former include Font et al. [11] and Porter et al. [50], while support

for the latter include Orlov [46], and Kim et al. [32]. Enloe et al. [8] however, found that

the momentum transfer during the positive-going half cycle was inconsistent in both

magnitude and direction, noting that competing force production may exist depending

on the details of the discharge structure in any given half cycle. The work of Enloe


et al. support the generally accepted view that the dominant force is provided by the

negative-going stroke of the AC cycle. Further discussion on the mechanisms of electro-

hydrodynamic force in plasma actuators extends beyond the scope of the current work,

thus the reader is referred to aforementioned literature on the subject.

2.2 |Operational Trends

Being fully electronic, plasma actuators are best characterized in terms of electrical quan-

tities. Currently, power consumed by the actuator, P , is typically characterized in terms

of power law relationships with the peak-to-peak voltage, Vpp, and frequency, f , of the

applied signal. Power as a function of voltage has been typically �tted to the curve

P = V npp with 2 < n < 3 [55, 45] or alternatively, P = (V − Vo)n with 2 < n < 3 [12]. A

number of groups have found P ∝ V7/2pp [9, 34]. Several power consumption and voltage

trends are summarized in Figure 2.3, adapted from Kriegseis [34].

pow

er c

onsu

mption

, P

[W

]

voltage, Vpp [kV]

2 3 4 5 10 20 30

10

10

10

10

2

1

0

-1

Figure 2.3: Power consumption as a function of applied voltage for various studies.

Kriegseis et al. [34] also found P ∝ f 3/2 for power dissipated as a function of applied

frequency contrary to previous publications which declared this a linear relationship

[12, 48, 50]. The declaration that P ∝ f would require that the energy consumed

per cycle is consistent over all frequencies, an oversimpli�cation which should become

apparent in Section 5.2.

Plasma actuators are also often characterized by the momentum transfer to neutral

air. This can be measured in terms of induced velocities or in terms of measurements of


Table 2.1: Induced velocity and applied voltage relation for plasma actuators.

umax experiment details

∝ V7/2pp PIV and LDV experiments

Enloe et al. (2004), Post (2004)∝ Vpp experiment and simulation

Orlov et al. (2006)asymptotically with Vpp Pitot tube experiments

Forte et al.(2007)∝ V n

pp with LDV experiments with MEMS actuators,n = 1.8, 1.9, 2.3 grounded electrode width = 10, 4, 1mm

at height of 0.5mm, Okochi et al. (2009)∝ V 2.4&2.7

pp PIV experimentvoltages below ≈ 10 kV Murphy et al. (2013)

the thrust exerted by the plasma actuator. The velocities induced by plasma actuators in

response to applied voltage have been modeled and measured with great variety amongst

studies. Reasons for discrepancies between results may include di�erences in actuator ge-

ometries, actuator materials, range of applied signals, measurement techniques, and high

voltage ampli�ers. Popular methods to obtain induced velocities are Pitot tube, par-

ticle image velocimetry (PIV), and laser Doppler velocimetry (LDV). Several maximum

induced velocity, umax, and voltage relations are summarized in Table 2.1.

2.3 |DBD Actuator Optimization

A number of works have focused on increasing the e�ectiveness of plasma actuators such

as the optimization studies done by Enloe et al. [9], Forte et al. [12], and Post [51].

In such studies, the e�ects of the applied AC signal, actuator geometry, and dielectric

thickness on the performance of plasma actuators have been considered.

Ionization occurs when the di�erence between the instantaneous AC potential and

the charge build-up on the dielectric surface exceeds a threshold value. This indicates

that there are optimal AC waveforms for actuator performance [4]. The square wave

is least e�ective, the sine wave is better, but the triangle wave is most e�ective [4, 40].

A waveform which extends the time allotted to the half cycle which provides the domi-

nant force and reduces the time allotted to the weaker half cycle, such as the sawtooth

waveform has been shown experimentally optimal [9]. Some studies have also shown that

there is an optimal frequency to achieve maximum induced velocity for a given actuator

geometry [45, 46].


The width of the encapsulated electrode has been shown to have a signi�cant e�ect

on the discharge of the plasma actuator. This buried electrode must be wide enough

such that the full extent of plasma can be generated. However, there is a critical width

at which the maximum injected velocity occurs, beyond which increasing the width of

the buried electrode is essentially ine�ectual [12]. It has been demonstrated that the

inclusion of a gap between the electrodes does not directly a�ect the performance of the

actuator but a slight overlap has a tendency to create a more uniform discharge ignition

[46]. Forte et al. [12] found that overlapping electrodes as well as electrodes with a large

separation show a decrease in induced velocity indicating the existence of an optimum

gap width.

Enloe et al. [9] concluded that changes to the dimensions of the exposed electrode

left the majority of the discharge characteristics unaltered but had a signi�cant e�ect

on the actuator performance. It was found that a thinner electrode resulted in greater

momentum transfer. Using round exposed electrodes, Hoskinson and Hershkowitz [24]

found a faster-than-linear trend for the increase of force as the exposed electrode diameter

was decreased.

The thickness of the dielectric is also an issue of consideration in plasma actuator

optimization. Thicker dielectrics can withstand higher voltages prior to material break-

down, allowing larger electric �elds and associated body forces [40]. More uniform plasma

discharge with fewer �laments is also associated with thicker dielectrics [46]. Recent evi-

dence presented by Thomas et al. [62] shows it may be advantageous to consider thicker

dielectrics with lower dielectric coe�cients to reduce power loss through the material,

allowing higher operating voltages and in turn higher body forces. In a comparison of

glass and polymethyl methacrylate (PMMA) dielectrics, Forte et al. [12] found that

glass allowed for larger electric �elds and resulted in higher maximum induced veloci-

ties when compared to PMMA, due to the higher dielectric constant of glass. At high

voltages however, the increased intensity of the electric �eld causes discharge to become

�lamentary and unstable. Power consumption was also greater in glass than in PMMA.

In a comparison of other dielectric materials, Te�on o�ered very similar results to glass

although at the same thickness glass produced more thrust with increasing voltage than

Te�on [5]. For low voltage ranges, thinning the dielectric has been shown to improve

achievable velocity [12] and e�ciency [45]. However, thin dielectrics risk increased �la-

mentary discharge resulting in less consistent measurement and material damage due to

local heating.


2.4 |Applications to Flow Control

Dielectric barrier discharges were �rst reported experimentally in 1857 with the ozone dis-

charge tube of W. Siemens. Since that time, dielectric barrier discharges have developed

applications in a wide variety of industries with applications such as ozone generation,

air puri�ers, �uorescent lamps, surface treatment, and display panels to name a few.

A detailed history of DBD applications can be found in [33]. DBD-based applications

have also gained popularity in the area of �ow control. Studies have demonstrated the

practical use of these actuators in separation control [55, 47, 30], aircraft noise reduction

[26, 37, 63], reducing losses in compressor blades [38], wake control [63], and boundary

layer control [15, 20, 19]. For �ow control applications, materials and geometries that

maximize induced velocity and increase the longevity of the actuator, as well as methods

which optimize precision of fabricated actuators are sought. These major considerations

led to the development of the microfabrication technique for plasma actuator manufac-

turing described in the present work.


Manufacturing

3.1 |Conventional Methods & Materials

Despite numerous laboratory examples, broader applications of plasma actuators have

been restricted by a number of practical limitations. For example, plasma actuators are

often constructed of self-adhesive copper foil tape a�xed to a dielectric material. Di-

electric materials used in plasma actuator research include polyimide [47, 9], quartz [63],

Te�on [40], ceramics [40], acrylic [49, 7], glass [45, 44], and FR-4 (woven glass and epoxy

resin laminate) [26, 56]. However, layers of polyimide (Kapton) tape are the most com-

monly used dielectric. While the vast majority of the plasma actuators reported in

literature have copper electrodes, notable exceptions include tungsten [24], stainless steel

[24], chromium/gold/chromium [45] and gold coated copper [27] electrodes. Hoskinson

et al. [24] demonstrated that induced forces are essentially independent of electrode

material, although dependent on exposed electrode thickness. The method of adhering

glue-backed metal �lm electrodes onto a dielectric material provides simple and inex-

pensive implementation of actuators on a variety of geometries, including both �at and

curved surfaces. However, this technique and most commonly used materials (copper

and Kapton tape) present signi�cant limitations to the reproducibility and durability of

the constructed actuators.

Typically, copper tape electrodes are cut and laid by hand. Due to the hand-made

construction, these devices su�er geometrical imperfections that are detrimental to pre-

cise and repeatable experimentation. Sharp points and wrinkles in the metal surface

create localized increases in charge concentrations which can lead to arcing through the

dielectric material [9]. Electrode thickness and shape are restricted by the discrete thick-

ness of the commercially available metal �lm tape. In addition, complicated electrode

geometries are di�cult to obtain by hand without major imperfections.

9

Chapter 3. DBD Plasma Actuator Manufacturing 10

Using polymer tapes, such as Kapton tape, as a dielectric also presents issues due

to discrete thickness, handling di�culty, and most signi�cantly, material degradation.

Polymers have poor resistance to the plasma environment and erode during extended

actuation periods. Degradation e�ects have been noted in plasma actuator studies using

polymer dielectrics such as PMMA and polyvinyl chloride (PVC) by Pons et al. [49].

Woven glass and epoxy resin laminates have been used in a number of studies, taking

advantage of printed circuit board (PCB) manufacturing techniques. Photolithography

is used to pattern copper clad boards and chemical etching removes unwanted metal

in the desired pattern. This technique allows for complicated geometries such as the

horseshoe and sinusoidal electrode geometries explored by Roy and Wang [56]. However,

PCB-based plasma actuators also degrade over time [52, 18]. In a similar approach, using

photolithographic and etching techniques, Durscher and Roy [7] were able to study horse-

shoe and sinusoidal electrode geometries by etching copper tape adhered to a dielectric

substrate of choice.

Okochi et al. [45] established a MEMS technique for the millimeter-scale fabrication

of plasma actuators. Using photolithography, 300 nm thick chromium/gold/chromium

electrodes were vapor deposited onto 0.525mm Pyrex glass substrates. Arrays of actua-

tors fabricated by Okochi et al. [44] were used to study the control e�ect in a turbulent

�ow at low to moderate Reynolds numbers.

3.2 |Device Degradation

Kapton and copper tape-based actuators are predominant in the �eld or plasma actuator

research due to experimental simplicity, i.e. readily available and inexpensive materials.

However, resultant actuator performance is limited by construction imperfections as well

as signi�cant material wear. As previously mentioned, hand-cut and laid metal foil

electrodes su�er from wrinkles and sharp points (present at both the macroscopic and

microscopic levels) which generate localized charge concentrations. Plasma discharge

is intensi�ed at these locations accelerating dielectric breakdown and actuator failure.

These imperfections are also unattractive from a �ow control perspective, as regions of

greater discharge intensities can cause irregularities in the induced �ow.

The degradation of polymers in plasma discharge is a known phenomenon [39, 64].

Polymer-based dielectrics fail to withstand the bombardment of ions, radical species

and UV radiation of the plasma environment and degrade during extended periods of

exposure [49, 52, 18, 13, 10] causing the operational properties and performance of ac-

tuators to change over time [18]. Furthermore, in the case of hand-layered polymer-tape


dielectrics, microscopic and visible air pockets between layers are inevitable, and the

potential for partial discharges between layers may also contribute to dielectric break-

down [42]. However, polymer-based dielectric materials and metal foils tapes for plasma

actuators remain widely used and little research has been conducted to examine the

e�ects of plasma-induced degradation on the operation of these actuators.

Layered Kapton tape dielectric, for instance, is comprised of alternating layers of poly-

imide �lm and silicone-based adhesive. With actuator usage, the top layer of polyimide

�lm degrades in the plasma-forming region adjacent to the exposed electrode to reveal

the foremost layer of adhesive [18]. The operational and performance characteristics of

an actuator therefore change during this process. The few studies which have noted the

degradation of polymer dielectrics such as PMMA and PVC [49], Kapton [10], and FR-4

[52], have been limited to visual interpretations. Recently, Hanson et al. [18] presented

quantitative information on the e�ects of plasma-induced dielectric degradation of plasma

actuators with commonly used dielectrics, Kapton and FR-4. The consumed electrical

power and actuator capacitance were used to monitor the health of the actuators over

prolonged actuation periods. Hanson et al. [18] included visual monitoring of the actu-

ator tests via a charge-coupled device (CCD) camera images to accompany operational

analysis. Variations in power consumption and e�ective capacitance of the actuator were

correlated to the visual documentation of physical changes to the actuator. Actuators

with Kapton dielectric exhibited degradation of both the dielectric material and the cop-

per electrodes resulting in increased power consumption and e�ective capacitance over

time. These electrical characteristics increased dramatically during the initial hours of

operation associated with the degradation of the top polyimide layer. The polyimide

layer degraded initially from the exposed electrode edge towards the streamwise extent

of the plasma-forming region. As the top polyimide layer degraded, the operational prop-

erties exhibited asymptotic behaviour. Changes to the operation of the actuator were

minimal beyond this initial degradation phase. These trends were more severe for both

higher operating voltages and frequencies, as well as for dielectrics comprised of fewer

Kapton layers.

3.3 |Methods for Increased Device Longevity

The degradation issues outlined in Section 3.2 are of consequence for the operation of

plasma actuators. As such, mitigation methods are required for both precise experi-

mentation as well as for the future applications of plasma actuators in industry. In

the current work, two methodologies are presented to enhance the longevity of plasma


actuators subjected to extended operation. Firstly, a protective surface treatment for

actuators comprised of non-degradation resistant materials is described in Section 3.3.1.

Secondly, a microfabrication process through which metallic �lm is precisely physical

vapour deposited onto glass substrates is described in Section 3.3.2.

3.3.1 |Protective Coating

A suitable protective coating was sought to increase stability of plasma actuators without

signi�cant reduction in performance. The aforementioned work of Hanson et al.. [18]

demonstrated that the silicone-based adhesive remains essentially intact on the dielectric

surface of an actuator following an extended period of actuation despite the degradation

of the polyimide layer. This result should not elicit surprise as silicone has long been

used for its electrical and heat insulating properties. Silicone-rubber has previously been

shown as a superior insulator against electrical discharges [60, 39]. Even the application

of silicone varnish has been shown to extend the lifetime of insulating materials exposed

to corona discharge [6].

Polydimethylsiloxane (PDMS) is a silicone-based product that is often used as an

insulator for high voltage applications due to its dielectric properties, low reactivity and

combustibility, and its high thermal and oxidative stability. Hillborg [22] found evidence

of thin glassy silica-like layer formation on the surface of a PDMS exposed to plasma

in air. This silica-like layer would have superior resistance to plasma due to its high

silicon and low organic content. Thus, PDMS was chosen for surface treatment of select

Kapton-based actuators in the current work with the idea that PDMS could provide

suitable protection of the actuator surface from plasma discharge. To allow for minimal

treatment thickness, a PDMS oil was selected. Each of the treated actuators tested in the

subsequent chapters received an application of Dow Corning PDMS oil (CSt 200) across

the entire actuator surface (including exposed electrode), followed by light polishing.

This surface treatment is a simple and inexpensive method aimed at increasing the usable

lifetime of degradation prone dielectrics, such as polymer tapes like Kapton.

3.3.2 |Microfabrication

In the current work, limitations of conventional construction methods are also addressed

by the development of a rigorous microfabrication procedure for plasma actuators with

an emphasis on selecting materials that can withstand the plasma environment. This

method confronts both construction precision and material degradation issues. An im-

portant objective in the process development was the ability to produce electrodes in


the micron thick range, as thinner exposed electrodes have been shown to impart greater

momentum transfer [9]. Using photolithography, a thin metallic �lm can be deposited

onto a dielectric substrate. Schott AF-45 alkali-free borosilicate glass was chosen as a di-

electric for reasons of compatibility with microfabrication processing, superior resistance

to material degradation in plasma, and dielectric properties preferable to basic borosil-

icate glasses, such as Pyrex. Polyimide �lm is not a viable dielectric option due to the

signi�cant thermal stresses during metal deposition, which lead to electrode failure, as

well as the aforementioned degradation issues.

1)

2)

3)

4)

5)

6)photoresistdielectric mask

lift-off resistmetal

UV light

Figure 3.1: Photolithographic process forplasma actuators, 2D cross-sectional view.

0.5 mm

Figure 3.2: The top-view of the undercutcreated with a bilayer resist process aftermetal deposition as seen through a micro-scope.

A bilayer resist process as shown in Fig-

ure 3.1 was optimized for actuator manu-

facturing allowing for metal depositions of

a few microns. This bilayer resist process

provides an undercut in the physical mask

after the development stage between the

deposited metal and the substrate for high

resolution metal lift-o�. The light out-

lined region surrounding the electrode in

Figure 3.2 shows the bilayer undercut pro-

�le which allows for clean edge metalliza-

tion and discourages delamination of the

deposited electrodes. The manufacturing

procedure was developed at the Western

Nanofabrication Facility at Western Uni-

versity in London, Ontario using a sput-

tering physical vapor deposition technique

for metal deposition. The sputtering de-

position tool employs a high voltage across

separated electrodes in a chamber of noble

gas, in this case argon, evacuated to low

pressure. The resulting electric �eld gen-

erates a glow discharge at the surface of

the source material. Argon ions bombard the surface of the source, ejecting material.

This metal vapor condenses on the substrate to form a thin �lm.

The most frequent cause of failure throughout the process derives from inadequate

cleaning of the substrate, leading to delamination of the metallic electrodes from the

glass surface. Small particles on the glass surface result in holes in the metal �lm surface.


Insu�cient rinsing of the developer chemical can cause entire electrodes to delaminate

during or after the lift-o� stage. When considering �lms in the micron-thick scale, a cause

of delamination is the increasing residual strain inherent to thin �lms with increasing

thickness. Direct deposition of copper onto glass is di�cult under ideal conditions, leading

to high percentage of failures. Thus, a titanium seeding layer was used to enhance copper

adhesion to the glass dielectric. Actuators fabricated with a seeding layer of titanium

on the order of hundreds of nanometers have a much higher success rate throughout

processing. Electrodes of tungsten, however, can be sputtered without a seeding layer

requirement with no adhesion issues.

Table 3.1: Spin-coat recipe speci�cations.

MaterialSpread Spin Bake

time [s] rpm time [s] rpm time [min] temperature [oC]LOR30B 7 500 45 2000 10 170S1827 5 500 45 2500 3 113PMMA 5 500 45 3000 3 180

The manufacturing procedure developed by the Flow Control and Experimental Tur-

bulence (FCET) lab is as follows [25]. Glass substrates are thoroughly rinsed with deion-

ized water to remove large scale contaminants. To avoid aforementioned delamination

issues, the glass is subjected to a twenty minute reactive ion etch process, where oxygen

plasma etches the surface of the glass eliminating surface contaminants including organic

material. A bilayer sequence of lift-o� resist (MicroChem LOR 30B) followed by positive

photoresist (MICROPOSIT S1827) are then spin-coated onto the glass substrate. Spin

recipes and bake requirements for the resists are shown in Table 3.1. The photoresist

layer is patterned via UV light exposure in a mask alignment tool. For the large features

of these devices, acetate masks provide su�cient resolution. UV activated photoresist

becomes soluble in photolithographic developer (MICROPOSIT MF-319) to expose the

substrate surface. A �ve minute ion etch is used to ensure cleanliness of the exposed

substrate prior to metal deposition. Metal is vapor deposited onto the substrate with a

sputtering deposition tool under an argon plasma. Organic solvent (MicroChem Nano

Remover PG) is used to lift-o� the metal deposited on the remaining photoresist while

metal deposited directly onto the glass substrate remains intact. A protective layer of

PMMA is spin-coated onto the freshly deposited electrodes such that the process can

be repeated on the virgin face of the substrate. Various stages of the microfabrication

process may be found in Figure 3.3. Low stress, homogeneous metallic �lms with low re-


sistivity are obtained using this method. Electrode metal can be sputtered to a thickness

of approximately 1µm. This procedure has been successfully applied to simple actuator

geometries up to a length of 7 cm on round glass wafers, as well as more complicated

arrays of several actuators up to lengths of 3 cm on square glass tiles.

(a)

(b)

(c)

(d)

(e) (f)

(g)

Figure 3.3: Various stages in microfabrication procedure: (a) spin-coat application ofLOR 30B, (b) the mask alignment tool used for UV exposure, (c) substrate preparedfor second side metal deposition, (d) four substrates loaded in the sputtering chamberfollowing metal deposition, (e) sputtering vapour deposition interface, (f) metal lifting-o�from glass substrate in solvent to reveal plasma actuator array, and (g) a microfabricatedplasma actuator device.

Photolithography is a costly and intricate procedure; however, the resulting actuator

constructions are of accuracy and repeatability unmatched by handmade counterparts.

Handmade plasma actuators can be assembled by a skilled craftsman with tolerances

reaching an upper limit of approximately ± 0.25mm for simple geometries. However,

devices fabricated with the photolithographic manufacturing process can be produced in

batches of several actuators with tolerances on the order of microns regardless of pattern

geometry.

As previously mentioned, the grounded electrode is typically encapsulated to eliminate

the occurrence of discharge on the underside of the actuator. This reverse discharge is

undesirable as a source of ine�ciency. Typically, Kapton actuators are laminated directly

onto non-conductive surfaces or built on top of additional layers of Kapton as a means

of encapsulation. However, this can allow for partial discharges at the unavoidable air

gaps between tape layers and electrode-Kapton junctions. Initial attempts to use Kapton

as an encapsulating material for a glass microfabricated actuator with an interelectrode


gap of 1mm demonstrated signi�cant reverse discharges within the insulating material.

exposed electrode

grounded electrode

(a) (b)

Figure 3.4: A 1mm interelectrode gap of a glassactuator exhibiting (a) partial discharges in Kap-ton encapsulation layers during actuation and (b)the resulting damage in insulating material.

Figure 3.4 clearly illustrates these �l-

amentary discharges and the con-

sequential pathways of damage in

the Kapton material. These branch-

ing pathways continued to propa-

gate with continued actuation and

degrade the polyimide material, fur-

ther promoting streamer formation.

To avoid the potential for partial dis-

charges, glass-based devices in the

current work were encapsulated with

glass bonder epoxy (Loctite E-30CL).

4 | Experimental Methods

For the experiments presented in the current work, the high voltage AC signal supplied

to plasma actuators was generated using a TREK20/20C high voltage ampli�er provided

oscillating signal from an Agilent 33210A waveform generator. The supplied voltage sig-

nal and voltage probe signals were monitored with a RigolDS1052E digital oscilloscope.

These pieces of equipment are shown in Figure 4.1.

plasmaactuator

voltage probecapacitor

high voltage amplifier

waveform generator

digital oscilloscope

microscope

CCD camera

Vpp

f = 1/period

{

Figure 4.1: Schematic of experimental set-up for degradation studies with CCD camera.

4.1 |CCD Camera Experimental Set-up

The degradation studies for which visual analysis is included in Chapter 6 were recorded

via StingrayF125B CCD camera mounted to an optical microscope, as shown in Fig-

ure 4.1. The images were recorded in low signal-to-noise ratio mode which averages 8

sequential images taken at a rate of 1.875 frames per second. One averaged image was

taken at the start of the degradation experiment and recorded every 3minutes thereafter.

The experiments were recorded in a lit laboratory space such that the actuator surface

17

Chapter 4. Experimental Methods 18

could be clearly viewed. As the light emitted by a plasma actuator is predominantly in

the ultra-violet region of the light spectrum [23], the CCD camera was essentially unable

to detect the plasma emission. Thus, the view of the actuator was not obscured by the

presence of plasma discharge.

4.2 |Electrical Characterization

In the current work, plasma actuators are characterized by their electrical properties,

namely their power consumption, P , cold capacitance, Co, and e�ective capacitance,

Ceff . Cold capacitance refers to the passive component of the actuator capacitance,

while e�ective capacitance refers the actuators capacitance with the presence of plasma.

Section 4.2.1 details the means by which these values are obtained via probe capacitor,

while Section 4.2.2 justi�es the choice of probe capacitor used in these experiments.

4.2.1 |Power Consumption & Capacitance Calculations

The power consumed by a plasma actuator is most commonly determined from recording

either the discharge current via probe resistor, or the charge via a probe capacitor.

Grundmann et al. [16] emphasize the convenience of the probe capacitor method, as the

capacitor integrates the current passing through the actuator in time, capturing all micro-

discharges. The probe voltage, Vp, measured across a probe capacitor, Cp, is proportional

to the charge, Q, crossing the electrodes, as expressed by

Q(t) = CpVp(t). (4.1)

The charge values can be plotted against the applied operating voltage to generate

Q− V cyclograms, also known as Lissajous �gures. The area inside the cyclogram repre-

sents the energy, Ek that is consumed per discharge cycle, k. Thus the energy consumed

can be calculated by

Ek =

∮k

Q(t)dV. (4.2)

Power consumed by the actuator per discharge cycle, Pk can then be determined using

the applied frequency, f by the expression

Pk = Ekf. (4.3)


Combining equations 4.1 to 4.3 and averaging over a total of K discharge cycles, leads

to the calculation of total power consumption, P , corresponding to the time traces of

V (t) and Vp(t) measured, written as

P = Ef =f

K

K∑k=1

∮k

CpVp(t)dV. (4.4)

An additional bene�t of the probe capacitor method is that capacitance information

can be extracted from the linear portions of the Q−V cyclogram [34, 35]. Both the cold

capacitance, Co, of the actuator and the e�ective capacitance, Ceff , which incorporates

the contribution to capacitance from the plasma discharge, can be determined from the

slopes indicated in Figure 4.2. The capacitances are typically taken as the average of two

slopes: one from the positive-going half cycle and one from the negative-going half cycle.

Since the cold capacitance represents the purely passive capacitance of the actuator, at

any given time these slopes can be expected to be nearly identical. As mentioned in

Section 2.1, the charge response of a plasma actuator is asymmetric between half cycles,

hence the notation C+eff and C−

eff for the slopes corresponding to the positive-going

and negative-going half cycles, respectively. To provide a single value of the e�ective

capacitance an average is typically reported, however, it is important to keep in mind

that this can mask signi�cant information about the characteristics of charge transfer

across electrodes. This point is illustrated in subsequent chapters.

200

150

100

50

0

-50

-100

-150

-200

-5 -2.5 0 2.5 5

char

ge, Q

[nC

]

applied voltage, V [kV]

Ceff

Co

sample dataaverage dataQ-V areacapacitances

Co

Ceff

Figure 4.2: A typical Q − V cyclogram fordetermination of electrical quantities.

For the present work, consumed elec-

trical power and capacitance quantities

were used to characterize devices and

monitor the health of these actuator over

prolonged actuation periods. These val-

ues were obtained by means of a probe

capacitor placed between the actuator

and ground as shown in Figure 4.1. Al-

though, cyclogram-based measurements

traditionally have a lower signal-to-noise

ratio than direct measurement of the

discharge current, post-processing is re-

quired to obtain robust information from

the cyclograms [34]. The mean charge

cycle in response to the applied signal

was found from raw data post-processed


with a Savitzky and Golay �lter [57] for the results presented in the current work.

4.2.2 |Probe Capacitor Independence Check

To obtain reliable measurements of the aforementioned electrical quantities with the

probe capacitor method, due diligence must be taken to select the appropriate capacitor.

This can be illustrated by considering the following relation,

1

Ctotal

=1

Cactuator

+1

Cp

. (4.5)

As Cp decreases, Vp increases, which increases the signal-to-noise ratio. However, since

the probe capacitor is connected in series with the actuator, the presence of the probe ca-

pacitor cannot be neglected for small Cp. In general, Cp is negligible for Cp >> Ceff > Co,

and it is important to validate the choice of capacitor satis�es that requirement while ob-

taining optimal signal-to-noise ratio [34]. The present results were obtained using a probe

capacitance of Cp=33nF, unless otherwise stated, consistent with the previous degrada-

tion studies by Hanson et al. [18]. This capacitance is three orders of magnitude greater

than Co and two orders greater than Ceff . Figure 4.3 demonstrates that Cp=33nF is

an appropriate selection for the microfabricated glass-based actuators featured in the

present work.

100 101 102

pow

er c

onsu

mption

, P

[W

]

probe capacitance, Cp [nF]

electrical circuit affected valuable results poor SNR

different capactiors

P = 0.53 W (at Cp = 33 nF)

0.6

0.55

0.5

0.45

Cp = 33 nF

Cp = 71 nF

Figure 4.3: Probe capacitor independence check for a microfabricated copper on glassactuator operated at 6 kV, 4 kHz.


4.3 |Particle Image Velocimetry

There are a number of ways the velocities induced by plasma actuators can be obtained.

Unlike Pitot tube measurements for instance, particle image velocimetry is preferable as

a non-intrusive means to obtain velocity information. This method has the additional

bene�t of acquiring the velocity components over the entire �eld of view (FOV) instan-

taneously. The experimental set-up used for all PIV experiments and the methods for

data processing are described in Section 4.3.1 and Section 4.3.2, respectively.

4.3.1 |PIV Experimental Set-up

Velocity components were acquired via a LaVision 2D PIV system. The PIV system

consisted of one Imager SX camera with a resolution of 2368 x 1776 pixels and a 105mm

focal length lens. Particles were illuminated with an Evergreen Nd:Yag laser with a

wavelength of 532 nm and pulse energy of 200mJ. A set of optics were used to create a

vertical, columnar light sheet, approximately 1.0mm thick. This optics sequence included

a cylindrical, divergent lens with radius 13.1mm that spread the laser beam, a convex

lens of radius 258.5mm to columnate the fanned laser light, and a spherical focusing

lens to focus the laser sheet. Laser sheet was oriented perpendicular to the actuator

surface and positioned in the approximate centre of the exposed electrode to minimize

the in�uence that 3D e�ects at the electrode ends may have. Tests were preformed in

a 0.58mx 0.56mx 0.58m acrylic box to ensure quiescent conditions during PIV image

acquisition. The quiescent box was seeded using an in-house Laskin-type seeder using

Di-Ethyl-Hexyl-Sebacat (DEHS) oil. Corresponding power measurements were acquired

within the image capture time frame (approximately 3.5 minutes for 1000 image pairs).

field of view

VpCp

Nd:Yag laser

PIV camera

Vac

Vpp

(a) (b)

x

z

y

Figure 4.4: (a)Schematic of experimental set-up and (b) actual set-up for PIV experi-ments with laser �ring.


4.3.2 |Data Processing

The raw PIV data was processed using LaVision's Davis 8.1.3 software. Data was pro-

cessed with an iterative multi-grid approach. The data was initially processed on a 32 x 32

pixel grid followed by a 16 x 16 pixel grid for �nal iterations with 50% overlap. A total

of 3 iterations were preformed on each image pair. A �eld of view of 28 x 21mm was

examined with a spatial resolution of approximately 0.09 x 0.09mm for the treated and

untreated Kapton-based actuator tests. A �eld of view of 32 x 24mm with associated spa-

tial resolution of approximately 0.11 x 0.11mm was recorded for all tests with glass-based

devices. The umax velocities discussed in Section 5.3 represent the maximum x-direction

velocity of the mean �ow �eld averaged from 1000 sequential snapshots of the velocity

�eld.

4.4 |Electromagnetic Field Measurement Set-up

The plasma actuator used in EMF experiments consisted of 35µm thick copper tape elec-

trodes adhered to 1.6 mm thick acrylic (PMMA). The grounded electrode was insulated

with 0.27mm of Kapton tape. The actuator had exposed and grounded electrode widths

of 0.5mm and 20mm, respectively, and an active length of approximately 120mm.

x

z

y

VpCp VacV

copper shield fieldmeter

height ladder

plasma actuator

(at 0° orientat ion)

plat form

measurement plane orientat ions:

plasma90°

0°

45°

315°270°

Figure 4.5: Schematic of experimental set-up for electric and magnetic �eld studies.


The actuator surface was oriented normal to a mock-�oor platform surface, as shown

in Figure 4.5 and able to pivot about its centre. Speci�c distances from the actuators

centre were indicated along the length of the platform. A wooden height ladder was

constructed such that measurements could be made at speci�c heights from the actuator

centre with the hand-held measurement device. Using the height ladder at the distances

indicated on the platform, the �eld measurements were obtained for a spatial grid of

points. The actuator was pivoted about its centre such that this spatial grid of �eld

measurements was obtained at various orientations relative to the actuator as de�ned in

Figure 4.5.

Both electric �eld and magnetic �eld strengths were determined using a hand-held

ME3951A GIGAHERTZ SOLUTIONS �eld meter. For repeatable electric �eld mea-

surements, the �eld meter was grounded and held in front of a square grounded copper

shield with dimensions of 50 cm as per the ME3951A manual. The �eld meter indicates

the RMS value of the �eld along the axis parallel to the device for electric �elds and per-

pendicular to the display screen for magnetic �elds. Each �eld of interest was measured

along three axes, as indicated in Figure 4.5, de�ned such that the y-axis is pointing in

the direction of the actuator center regardless of measurement plane orientation. The

resultant of three-axis measurements was recorded for each height, distance, and orien-

tation in order to quantify the magnitude of the �eld strength at each of these locations

surrounding an operating plasma actuator.

5 | Characterization of

Microfabricated Devices

5.1 |Actuator Speci�cations

groundedelectrode

d

L

we

de

dg

wg

topview:

cross-section: memd

Figure 5.1: Plasma actuator dimension de�-nitions.

Figure 5.1 illustrates the geometric and

construction de�nitions of plasma actu-

ators that will be referred to in subse-

quent sections. These de�nitions include

the electrode material, me, dielectric ma-

terial, md, dielectric thickness, d, exposed

electrode thickness, de, grounded electrode

thickness, dg, exposed electrode width, we,

grounded electrode width, wg, and active

length of the actuator, L. The active

length, L, excludes the rounded ends of

the electrodes which are present to reduce

sharp points and localized charge build-up.

In the current chapter, microfabricated plasma actuators are examined with respect to

electrical behaviour and induced velocities. The induced velocity pro�les of glass-based

actuators, a PDMS treated Kapton actuator, and an untreated Kapton actuator are

compared. The glass-based plasma actuators used in the current chapter were constructed

using the microfabrication procedure outlined in Section 3.3.2 while tape-based devices

were constructed by hand. Consistent dimensions of we=5mm, and wg=10mm were used

for these devices. The actuators are summarized in Table 5.1.

24

Chapter 5. Characterization of Microfabricated Devices 25

Table 5.1: Summary of plasma actuators used in characterization of microfabricateddevices. Open symbols indicate PIV experiments, ie) seeded conditions, whereas solidsymbols indicate an environment free of seeding.

symbol L [mm] md me d [mm] de [µm] dg [µm]� 200 Kapton Cu tape 0.18 35 35© 200 Kapton (PDMS) Cu tape 0.18 35 35

�,�,H, ©, B, ♦ 80 AF-45 Cu 0.3 1.0 1.0♦,�,O, ©, B 80 AF-45 W 0.3 1.0 ∼0.5

5.2 |Electrical Quanti�cation

The variations in power consumption per unit length with applied frequency are shown

for microfabricated actuators with both copper and tungsten electrodes are shown in

Figure 5.2a. Results plotted with solid symbols were obtained with a probe capacitance,

Cp=71nF, while all other tests were conducted with Cp=33nF. For actuators in a seeded

environment (open symbols), the power consumption demonstrates a weak power law

dependence with frequency, P/L ∝ f 1.1±0.3. This agrees with the P/L ∝ f 1.15 depen-

dence from Hanson et al. [18] for Kapton-based actuators of similar dielectric thickness

(d =0.36mm). The power law dependence supports the claim by Kriegseis et al. [35]

that the energy consumed per discharge cycle is a�ected by changing frequency and the

linear relation reported in a large number of other studies is, in fact, an oversimpli�-

cation. However, because the power law is subtle in nature, one could argue that the

results shown here lie within experimental uncertainty of the linear evolution of power

consumption with applied frequency commonly reported by other researchers [12, 48, 50].

A linear evolution would require that the energy consumed per discharge cycle is constant

regardless of applied frequency, which is not the case for the majority of cases shown here.

Forte et al. [12] demonstrated that the slopes of power-frequency relations increased

with increasing applied voltage. This e�ect is also observed in the current work for the

two unseeded copper on glass actuator cases at 6 kV (�) and 8 kV(�). A notable di�erence

in slope is also observed between seeded and unseeded copper on glass cases at the same

operating voltage and frequency. This discrepancy can be explained by the di�erence in

probe capacitors used to obtain the results. As illustrated in Figure 4.3, changing the

probe capacitance can have a signi�cant e�ect on results, and using probe capacitors

of higher capacitance reduce the signal-to-noise ratio of acquired data. In addition, the

presence of seeding particles likely has a subtle e�ect on actuator operation. Support for

these suppositions lies in the similarity between measurements obtained with identical


probe capacitance for actuators in seeded conditions.

2 4 6 8 100

5

10

15

20

25

30

35

frequency, f [kHz]

P/L [W

/m]

4 6 8 10 120

10

20

30

40

50

60

70

80

voltage, Vpp [kV]

P/L [W

/m]

(a) (b)

Cu on glass

Vpp

Cu on glass

Vpp

W on glass

Vpp

4.0 0.3

3.8 0.3

3.3 0.4

f1.0 0.2

f1.1 0.3

f1.2 0.2

Cu on glass, 6 kV

Cu on glass, 8 kV

Cu on glass, 6 kV

W on glass, 6 kV

Figure 5.2: Power consumption per unit length as a function of (a) frequency and (b)applied voltage for microfabricated actuators (at 4 kHz).

The power consumption as a function of applied peak-to-peak voltage (Vpp) is shown

in Figure 5.2b. For actuators in a seeded environment, the power consumption demon-

strates a power law dependence with voltage, P/L ∝ V 3.8±0.3pp and V 3.3±0.4

pp for copper

and tungsten electrodes, respectively. These lie within the range of power law relations

P/L ∝ V npp where 2 < n < 3.5 commonly published in other studies [12, 18, 41], in-

cluding the prominent relation P/L ∝ V 3.5pp [9, 34]. Hanson et al.. reported the relation

P/L ∝ V 3.3pp for Kapton-based actuators of similar dielectric thickness (d =0.36mm) and

operating conditions. Similarities with the results of Hanson et al. are particularly rel-

evant to justify comparisons between microfabricated actuators and handmade Kapton

devices in Chapter 6. By using identical equipment to that of Hanson et al., the present

work veri�es that these microfabricated actuators operate according to expected power

law relations. These results are also similar to the P/L ∝ V 2.7pp relation reported by

Okochi et al. [45] for MEMS manufactured plasma actuators of Cr/Au/Cr electrodes on

0.5mm Pyrex glass.

Relative cold and e�ective capacitances are shown in Figure 5.3a as a function of

operating frequency and as a function of operating voltage in Figure 5.3b. The value of


(a) (b)

Co/L [pF/m

]

Co/L [pF/m

]Ceff/L [pF/m

]

Ceff/L [pF/m

]

frequency, f [kHz] voltage,Vpp [kV]

50

100

150

200

250

2 4 6 8250

300

350

400

450

500

550

50

100

150

200

250

4 6 8 10200

400

600

800

1000

Cu on glass, 6 kVCu on glass, 8 kVW on glass, 6 kV

Cu on glassW on glass

Figure 5.3: Cold and e�ective capacitances per unit length as a function of (a) frequencyand (b) applied voltage (at 4 kHz) for microfabricated actuators.

Co/L represents the purely passive component of the actuator and therefore is maintained

despite increasing frequency or voltage. The e�ective capacitance, however, includes the

contribution to capacitance from both the actuator and the plasma discharge. The

lower portion of Figure 5.3a demonstrates that Ceff/L can be expected to increase with

frequency in the cases where weak power law dependence was observed between power

consumption and frequency. Alternatively, for the unseeded copper on glass case at 6 kV,

which exhibited a linear relation between power consumption and frequency, the e�ective

capacitance per length is relatively constant. As the voltage applied to an actuator is

increased, the associated rise in electric �eld strength enlarges the streamwise extent of

the plasma forming region [9, 34]. The charged plasma region above the surface of the

grounded electrode virtually extends the width of the exposed electrode. For parallel

bodies of charge, the capacitance is proportional to the area of overlap. Therefore, in the

case of a plasma actuator the increasing area of virtual and grounded electrode overlap

with voltage causes the e�ective capacitance to also increase. This behaviour is shown

in the lower plot of Figure 5.3b.


−150

−100

−50

0

50

100

150

−3

−2

−1

0

1

2

3

charg

e [n

C] v

olta

ge [kV

]

−150

−100

−50

0

50

100

150

−3

−2

−1

0

1

2

3

charg

e [n

C] v

olta

ge [kV

]

−150

−100

−50

0

50

100

150

−3

−2

−1

0

1

2

3

cha

rge

[nC

] volta

ge [kV

]

−150

−100

−50

0

50

100

150

0.0 0.05 0.1 0.15 0.2 0.25

charg

e [n

C]

t ime [ms]

0.0 0.2 0.4 0.6 0.8 1.0

t reated KaptonCu on glassW on glass

(a)

(b)

(c)

(d)

Figure 5.4: Charge across the electrodes in re-sponse to applied voltage for (a) a treated Kap-ton actuator, (b) a copper on glass actuator, and(c) a tungsten on glass actuator with (d) themean cycle after 60 hours of continuous opera-tion at 6 kV, 4 kHz.

Comparatively, a higher degree of

variation was observed in the electri-

cal quantities of the actuators with

tungsten electrodes. This scatter in

electrical quantities is due to the na-

ture of the charge crossing the elec-

trodes and the method used to deter-

mine the electrical quantities. As out-

lined in Section 4.2, the slopes of the

linear portions of the Q − V cyclo-

grams represent the capacitances at

various points in the AC cycle. The

values previously quoted for Ceff are

the average of C+eff of the positive

half cycle and C−eff of the negative

half cycle. The same method applies

to the calculation of Co. However,

as discussed in Section 2.1, plasma

actuators have an asymmetrical cur-

rent (and therefore, charge) response

to an AC signal. For Kapton-based

actuators the di�erence is subtle, as

shown in Figure 5.4a. Alternatively,

Figures 5.4b and 5.4c demonstrate the

asymmetry of the charge response to

the applied signal, as well as the sig-

ni�cant spikes in charge crossing the

electrodes that occur for microfabri-

cated actuators with tungsten elec-

trodes. These large spikes distort

the shape of the average cycle. Fig-

ure 5.4d shows the mean charge cycles,

averaged from approximately 200 cy-

cles. This distortion is more severe

in the positive-going half cycle, as ex-

pected. Although present for both microfabricated actuators, the spikes in charge and


thus, the distortion of the averaged cycle is more severe with tungsten electrodes.

−3 −2 −1 0 1 2 3

+

_

+

_

150

100

50

0

-50

-100

-150

char

ge, Q

[nC

]signal voltage [kV]

Ceff

Co

Co

Ceff

treated KaptonCu on glassW on glass

Figure 5.5: Variations in a Q − V cyclogramfor microfabricated and handmade actuators af-ter 60 hours of operation at 6 kV, 4 kHz. Thesubstantial di�erence in cyclogram area betweenhandmade and microfabricated actuators is dueto di�erences in actuator length.

10 12 14 16 18 20

80

70

60

50

40

30

20

time [hrs]

Ceff

[pF]

Ceff, cyclogram

Ceff, cyclogram

Ceff, histogram

_

+

Figure 5.6: The e�ective capacitance values cal-culated from the positive half cycle slopes of cy-clograms (C+

eff ), the negative half cycle slopesof cyclograms (C−

eff ), and using the histogrammethod for a microfabricated tungsten on glassactuator over a period from 10 to 20 hours intocontinuous operation at 6 kV, 4 kHz.

Figure 5.5 further demonstrates

the e�ect the distortion of the mean

cycle has on the extraction of capac-

itance information. Clear linear por-

tions of the cyclograms corresponding

to the passive component of the ac-

tuator exist and agree with minimal

variation for all devices. For Kapton-

based actuators, C+eff and C

−eff can be

extracted without di�culty, whereas

the e�ective capacitance for micro-

fabricated actuators is obtained with

greater uncertainty. The linear slope

corresponding to the C−eff of the

negative-going half cycle is apparent

without confusion and o�ers values

without signi�cant scatter. During

the positive-going half cycle however,

the cyclogram features a staircase of

linear portions which represent C+eff

at various point in the cycle caused by

the distortion in the mean charge cycle

as discussed with Figure 5.4d. The ap-

proximately linear segment closest to

the maximum charge across the elec-

trodes provides C+eff values most sim-

ilar to C−eff however with greater as-

sociated uncertainty.

Figure 5.6 compares the e�ective

capacitances determined from cyclo-

grams as shown in Figure 5.5 with val-

ues for Ceff determined using a his-

togram method for a series of mea-

surements over a 10 hour period. The

histogram method extracts the capac-


itances which occur most frequently throughout the approximately 200 cycles sampled.

These capacitances are derived by means of forward di�erencing of the recorded charge

and voltage time traces according to

C(ti) =∆Q

∆V|ti =

Q(ti+1)−Q(ti)

V (ti+1)− V (ti). (5.1)

The results from the histogram method follow a similar trend to that of the results

extracted from the C−eff slope of the cyclogram. This demonstrates that although a

range of e�ective capacitances occur for microfabricated devices, the dominant e�ective

capacitance matches that extracted from the negative half cycle discharge. The values of

C−eff demonstrate that the nature of the negative-going discharges is more stable over the

same time period for which C+eff has signi�cant �uctuations and associated uncertainty.

Since C−eff values are typically found to be the most dominant capacitance value, one may

be tempted to consider only these values as the `true' e�ective capacitance. On the other

hand, the values of C+eff demonstrate that the nature of the positive-going discharge is

more irregular than the negative-going discharge for microfabricated actuators. As such,

the values of C−eff and C+

eff will both be shown in the degradation studies of Chapter 6.

Secondary electron emission, which plays a critical role in sustaining plasma discharge,

occurs when a material is bombarded with electrons, such as the exposed electrode during

the positive-going half cycle of the applied waveform. Properties of the plasma discharge

can be sensitive to the e�ective secondary electron emission coe�cient of the electrode

and dielectric surfaces involved [58]. Since the maximum secondary emission of tungsten

is greater than that of copper [1], the increased electron emission by tungsten may account

for the greater occurrence of spikes in charge crossing the electrodes during the positive-

going half cycle. Hoskinson et al. [24] reported that changing the secondary emission

coe�cient of the exposed electrode did not have a signi�cant e�ect on plasma actua-

tor force production; a �nding that is supported by the comparable induced velocities

for microfabricated actuators with di�erent electrode materials shown in the following

chapter.

5.3 |Momentum Transfer to Air

Velocity �elds were obtained via PIV as described in Section 4.3 such that maximum

induced velocities could be extracted. Maximum induced velocity as a function of power

consumption, applied frequency, and applied voltage is shown in Figure 5.7 for microfab-

ricated actuators. Actuators with copper and tungsten electrodes follow similar curves


within experimental uncertainty for all properties examined. The maximum induced ve-

locity increases according to similar power laws for power consumption and frequency.

Increase in power consumption, as a result of increasing operating voltage at 4 kHz, pro-

duced the relations umax ∝ P 0.74±0.08/L and umax ∝ P 0.75±0.09/L for copper and tungsten

electroded actuators, respectively. Jolibois and Moreau [29] found that the maximum

induced velocity increased asymptotically with power consumption for PMMA-based

actuators with dielectric thicknesses ranging from 0.5mm to 5mm. Although Jolibois

and Moreau found that an asymptote was reached beyond 200W/m, the velocities they

presented for the range of power consumption values of the current work match those

presented in Figure 5.7a. Induced maximum velocity as a function of operating frequency

at 6 kV can be expressed as umax ∝ f 0.8±0.2 and umax ∝ f 0.7±0.4 for copper and tungsten

electroded actuators, respectively.

0 20 40 60

2 4 6 8 10

4 6 8 10

0.75 0.09

Cu on glass

P

W on glass

P

0.74 0.08

0.7 0.4

Cu on glassfW on glassf

0.8 0.2

2.7 0.3

Cu on glass

Vpp

W on glass

Vpp

2.5 0.2

(a) (b) (c)

um

ax [m

/s]

um

ax [m

/s]

um

ax [m

/s]

4.5

4

3.5

3

2.5

2

1.5

1

0.5

0

2.4

2.2

2

1.8

1.6

1.4

1.2

1

0.8

0.6

0.4

4.5

4

3.5

3

2.5

2

1.5

1

0.5

0

P/L [W/m] frequency, f [kHz] voltage, Vpp [kV]

Figure 5.7: Maximum induced velocity as a function of (a) power consumption (withincreasing voltages at 4 kHz), (b) frequency (at 6 kV), and (c) applied voltage (at 4 kHz)formicrofabricated actuators.

For microfabricated actuators, maximum induced velocity scaled according to V 2.5±0.2pp

and V 2.7±0.3pp for copper and tungsten actuators, respectively. These results compare to

the �ndings of umax ∝ V 2.4&2.7pp published by Murphy et al.. [41] for copper tape, 0.36mm

Kapton-based actuators at 4 kHz and voltages below ≈ 10 kV. The maximum velocity

induced at a height of 0.5mm scaled with V 3.12±0.07pp for the microfabricated devices, a sig-

ni�cant increase in momentum transfer from the relation umax,y=0.5mm ∝ V 1.8pp reported by

Okochi et al. [45] for MEMS manufactured actuators (Cr/Au/Cr on Pyrex) with 10mm


wide grounded electrodes. Enloe et al. [9] reported that both umax and power consump-

tion scale according to V 3.5pp for Kapton and copper tape-based devices, indicating a linear

relation between power consumed and induced velocity. Whereas, from the previous sec-

tion microfabricated actuators were found to consume power proportional to V 3.3&3.8pp ,

thus a linear relation between power and velocity is not observed. An exponent less than

unity can be expected due to the fact that the discharge becomes �lamentary and un-

stable at higher operating voltages, which results in loss of momentum transfer [12, 40].

This accounts for the asymptotic behaviour noted by other studies [12, 29] for voltages

which result in power consumption beyond a certain threshold. According to Moreau [40],

the di�erence in asymptotic versus power law relationships between umax and Vpp are

explained by the di�erent power supplies, actuator geometries, and measurement tech-

niques. Most signi�cantly, the results presented in Figure 5.7 agree with power law de-

scriptions from other PIV studies using similar equipment with polymer-based dielectrics.

y/δ

3

2.5

2

1.5

1

0.5

00 0.2 0.4 0.6 0.8 1

Kapton, umax=0.64 m/s, δ1/2=0.49 mmtreated Kapton, umax=0.88 m/s, δ1/2=0.55 mmCu on glass, umax=0.76 m/s, δ1/2=0.47 mmW on glass, umax=0.74 m/s, δ1/2=0.52 mm

1/2

u/umax,10

Figure 5.8: Velocity pro�les, normalized according to Equa-tion 5.3, of �ow induced by a actuators of various construc-tion methods and materials,each operated at 7.5 kV, 4 kHz.

Conventionally, wall jet

velocity pro�les are charac-

terized by their thickness,

δ and maximum jet veloc-

ity, umax [14, 34]. Due to

di�culties associated with

accurate δ estimation from

experimental data, the wall

normal distance of the 50%

maximum jet velocity can

instead be used as a robust

spatial measure. In this

section, δ 12is used as a nor-

malization factor to scale

the velocity pro�les and is

de�ned,

δ 12

= y(umax

2). (5.2)

Figure 5.8 shows the wall jet pro�les for a microfabricated actuator with copper elec-

trodes, a microfabricated actuator with tungsten electrodes, a typical handmade Kap-

ton actuator, and a handmade Kapton actuator with the protective PDMS treatment

described in Section 3.3.1. These pro�les are located at x=10mm downstream of the


plasma-forming edge and normalized by,

y∗ =y

δ 12,10

, u∗ =u

umax,10

. (5.3)

Jukes et al. [31] and Murphy et al. [41] present the theoretical velocity pro�les normalized

in this manner for both laminar and turbulent wall jets, according to Glauert [14]. Based

on these theoretical pro�les, one can expect y∗(umax) ≈ 0.5 and y∗(umax) ≈ 0.25 for

laminar and turbulent wall jets, respectively. With this guideline is can be concluded that

all plasma actuators tested here induced laminar wall jets. Microfabricated actuators had

nearly identical umax normalization factors, and all actuators showed a 50% maximum

jet velocity height of approximately 0.5mm. The untreated Kapton actuator induced the

weakest wall jet of all constructions tested. All pro�les correspond to new actuators (run

time prior to measurement less than 5minutes) with similar thicknesses of 0.27mm and

0.3mm for Kapton-based and glass-based devices, respectively. Devices were operated

at 7.5 kV, 4 kHz for consistent volts per thickness ratios.

6 | Device Degradation Studies

6.1 |Actuator Speci�cations

The plasma actuators featured in the current chapter have the consistent dimensions of

we=5mm, and wg=10mm. Other device properties are summarized in Table 6.1. All of

the tests discussed in this chapter were conducted in unseeded environments.

Table 6.1: Summary of plasma actuators used in degradation studies recorded via over-head CCD camera during continuous operation at 6 kV, 4 kHz.

symbol L [mm] md me d [mm] de [µm] dg [µm]� 200 Kapton Cu tape 0.18 35 35© 200 Kapton (PDMS) Cu tape 0.18 35 354 200 Kapton (PDMS) Cu tape 0.27 35 35B 200 Kapton (2xPDMS) Cu tape 0.18 35 35♦ 80 AF-45 Cu 0.3 1 1O 80 AF-45 W 0.3 1 ∼0.5

6.2 |PMDS Treated Kapton Actuators

The e�ect of a protective PDMS surface treatment on hand-laid copper and Kapton

actuators was investigated for three main cases: a 0.18mm thick actuator with single

PDMS treatment (©), a 0.27mm thick actuator with single PDMS treatment (4), and a

0.18mm thick retreated actuator (B). All cases were treated as outlined in Section 3.3.1

prior to operation at 6 kV, 4 kHz. The retreated case was stopped momentarily after 25

hours of operation for a second PDMS application. All Kapton-based actuators were

constructed consistent with Hanson et al. [18]. The results for an untreated hand-laid,

0.18mm thick Kapton actuator (�) from [18] are included for reference.

The time, t, evolution of these actuators power consumption, P , e�ective capacitance,

Ceff , and cold capacitance, Co is shown in Figure 6.1. The changes in electrical charac-

34

Chapter 6. Device Degradation Studies 35

teristics were monitored over 60 hours and determined using the methods described in

Section 4.2. The standard deviation of the measured power per cycle was typically ≤4%for the Kapton-based actuators discussed in this present section.

P/P0

Ceff/Ceff,0

Co/Co,0

time [hrs]

0 10 20 30 40 50 60

1.4

1.2

1

0.8

1.4

1.2

1

0.8

1.05

1

0.95

0.9

0.85

(a)

(b)

(c)

Kapton(0.18 mm)treated Kapton(0.18 mm)treated Kapton(0.27 mm)retreated Kapton(0.18 mm)

Figure 6.1: Variations in the (a) power consumption, (b) e�ective capacitance, and (c)cold capacitance of various Kapton-based actuators during 60 hours of continuous oper-ation at 6 kV, 4 kHz. Values shown are normalized by respective initial measurements att=0hrs.

The increase in power consumed by the untreated Kapton actuator is shown in Fig-

ure 6.1a. The normalized power consumption increased almost linearly for t < 8 hrs at

a rate of approximately 0.04/hr. The power consumption asymptotically reached steady

state operation at a value 40% higher than the initial value at t=0hrs. The e�ective

capacitance of the untreated actuator, found in Figure 6.1b, also featured asymptotic

behaviour following a nearly linear increase of approximately 0.03/hr over the initial

8 hours of operation. The e�ective capacitance increased by 31% over a 44 hour period

of continuous operation. Cold capacitance demonstrated a minor decrease during the

period of steepest increase in P and Ceff , as shown in Figure 6.1c.

The PDMS preparation signi�cantly altered the typical degradation behaviour of


Kapton actuators described above. The 0.18mm and 0.27mm thick treated actuators

exhibited a reduction in power consumption to a minimum of 85% of the initial value,

which occurred at approximately 30 hours into operation. Both of these actuators expe-

rienced nearly linear increases in power consumption thereafter. After 60 hours of oper-

ation the 0.18mm thick case had reached a power consumption increase of 13%, while

the 0.27mm case �nished approximately at the starting value. The rates of increase in

normalized power consumption (calculated for t> 48 hours) were roughly 1.2%/hr and

0.6%/hr for 0.18mm and 0.27mm thick treated actuators, respectively. By increasing

the dielectric thickness by 50% the rate of power consumption increase is reduced by

a factor of 2. Similar trends were also observed for the 0.18mm and 0.27mm treated

Kapton actuators in the Ceff values, shown in Figure 6.1b. E�ective capacitance for

both actuators reached minimums of 85% of the initial values and exhibited nearly lin-

ear increases following 30 hours of operation. Cold capacitance changes were minor for

all PDMS treated cases with the exception of the 0.18mm single treatment case. For this

case, Co decreased linearly after 30 hours of operation to 90% of the initial value after

60 hours of operation. This e�ect was not observed in the 0.27mm treated case.

The �nal case was treated with PDMS prior to operation and again after t=25hrs.

This retreated case exhibited similar reduction in P and Ceff to the single treatment cases

for the hours following the second PDMS application after 25 hours of operation. The

power consumption and e�ective capacitance exhibited an initial spike in properties after

re-ignition of the plasma after 25 hours (actuator turned o� for reapplication) followed

by a reduction in these characteristics that was maintained for the remainder of the ex-

periment. Unlike the single treatment case of the same thickness, onset of linear changes

to actuator operation after 30 hours of operation were not observed. The retreated case

instead completed the 60 hour test with a 13% reduction in power consumption.

The degradation tests were also monitored visually for physical degradation via an

overhead CCD camera as outlined in Section 4.1. The corresponding plasma-induced

degradation of the actuator surface is shown in Figure 6.2. The destructive e�ects of the

plasma environment are most apparent along the edge of the exposed electrode at the

interface with the dielectric material upstream of the grounded electrode, referred to as

the plasma-forming edge.

The untreated 0.18mm thick Kapton actuator shown in Figure 6.2a exhibited severe

degradation of the plasma exposed polyimide dielectric surface. This degradation initi-

ated at the plasma-forming edge and continued towards the extent of the plasma-forming

region. Clouding of the dielectric surface was observed as early as 0.5 hours into oper-

ation. Over 44 hours of actuation, the entire thickness of the foremost polyimide layer


0 hr 12 hr 24 hr 36 hr 48 hr 60 hr

3.4

mm

dielectric

3.4

mm

0 hr 2 hr 4 hr 8 hr 16 hr 32 hr

0 hr 12 hr 24 hr 36 hr 48 hr 60 hr

3.4

mm

(a)

(b)

(c)

exposedelectrode

dielectric

exposedelectrode

dielectric

exposedelectrode

domain

Figure 6.2: CCD images of exposed electrode (top) and dielectric surface (bottom)for (a)hand-laid copper on 0.18mm Kapton (from [18]), (b)hand-laid copper on PDMStreated 0.18mm Kapton, and (c)hand-laid copper on twice PDMS treated 0.18mm Kap-ton actuators during continuous operation at 6 kV, 4 kHz.

appeared to degrade in the plasma-forming region to expose the silicone-based adhesive

underneath. This period of physical degradation corresponded to the increases in power

consumption and e�ective capacitance. The reduction in dielectric thickness results in a

greater voltage per Kapton thickness ratio, thus, decreasing resistance, R to charge trans-

fer and increasing power consumption (P = V 2/R). Furthermore, recall that C ∝ εr/d.

Therefore, as the dielectric thickness, d, between the grounded electrode and the virtual

electrode (plasma) decreased, the actuators e�ective capacitance increased. The asymp-

totic trends of the power consumption and e�ective capacitance can be attributed to the

nature of the dielectric degradation. The initial steep increase in electrical quantities

corresponds to the Kapton erosion along the plasma-forming edge to the depth of the

silicone adhesive which deterred further wall normal erosion. The rate of change for elec-


trical quantities decreased as the erosion continued in a streamwise direction and reached

steady values as a streamwise degradation extent was reached.

The cold capacitance, decreased slightly during this period as a substantial portion

of the dielectric had been replaced with air (εair < εKapton). This decrease in dielectric

constant lowers the capacitance of the actuator itself. The actuator cold capacitance may

also have decreased as a result of receding or lifting of the copper tape exposed electrode

and weakening of the induced electric �eld.

Figure 6.2b shows the time evolution of actuator surface images for the 0.18mm thick

treated actuator. With a single PDMS surface treatment, the Kapton dielectric sur-

face remained intact after 12 hours of operation. The clouding of the polyimide surface

in the plasma-forming region observed in the untreated case failed to appear in the

treated case. Typically, for untreated actuators, surface degradation initiates at the

plasma-forming edge and moves streamwise towards the extent of the plasma region.

In the oil treated cases, however, degradation occurs sporadically along the plasma-

forming edge forming coral-like degradation patterns which branch out radially. The

onset of substantial degradation patterns between 30 and 40 hours into operation co-

incide with the onset of changes in power consumption and e�ective capacitance ob-

served for treated actuators. Power consumption and e�ective capacitance increase in

a linear fashion as these degradation patterns develop in size and merge together. The

presence of the silicone oil coating delayed and reduced the slopes of linear increases

in electrical properties typically observed during initial erosion of a Kapton actuator.

1 mm

Figure 6.3: Surface of PDMS treated Kapton-based actuator (0.18mm thick) following 60hours of continuous operation at 6 kV, 4 kHz.

The retreated case, pictured in Fig-

ure 6.2c, showed minimal signs of degra-

dation occurring rarely as small irregu-

larly shaped holes in the dielectric sur-

face. The maintenance of electrical prop-

erties can be attributed to the lack of

physical degradation exhibited by re-

treated actuator.

In addition to the reduction in surface

degradation and maintenance of electrical

properties with oil application, of partic-

ular interest is the signi�cant alterations

to the degradation patterns. Although

large congregations of coral-like degrada-

tion patterns mar the surface of the treated actuators following heavy use, as shown in


Figure 6.2b, other areas can remain entirely free of these blemishes. A larger area of actu-

ator surface is shown in Figure 6.3, demonstrating the sporadic nature of the degradation

patterns on a PDMS treated actuator after 60 hours of operation.

Visual monitoring and scanning electron microscope (SEM) images revealed that

once small divots appeared in the Kapton surface, the degradation patterns propagated

and developed into irregular shapes over time. As these patterns increased in size and

occurrence, the patterns merged together forming larger degraded regions with irregular

edges. The degradation patterns appeared to initiate from weak spots in the dielectric

surface near the plasma-forming electrode edge. After heavy usage large degraded patches

occurred at the electrode interface and initiation spots began to occur at locations on

the dielectric surface close to the electrode edge, as seen in Figure 6.4a. These initiation

spots were found up to 0.5mm away from the electrode edge. For the actuator retreated

at t=25hrs, the majority of physical degradation occurred as round holes in areas away

from the plasma-forming electrode edge as shown in Figure 6.4b. These holes in the

dielectric were infrequent along the length of the actuator.

Energy-dispersive X-ray spectroscopy (EDS) analysis found that trace amounts of

elemental silicon occur on intact portions of the dielectric surface, as well as on the

surface of the exposed electrode. This may indicate that residual silicon from the oil

treatment is deposited on the actuator surface. In the degraded regions, prominently

silicon was found, indicative of complete Kapton degradation down to the silicone-based

adhesive layer.

plasma-forming edge

initialization holes

degraded area

initiation holes

plasma-forming edge

initiation holes

plasma-forming edge

(a) (b)

Figure 6.4: SEM images of (a) degradation patterns in the dielectric surface of siliconetreated 0.18mm Kapton actuator and (b) initiation holes in the dielectric surface of a0.18 mm Kapton actuator with two oil treatments. Both actuators were operated for60 hours at 6 kV, 4 kHz.

These results show the protective capacity of the PDMS oil application as well as


supports the hypothesis that vulnerable spots in the dielectric surface act as initiation

locations for the propagation of fringed degradation patterns.

6.3 |Microfabricated Glass Actuators

The e�ects of long-term operation were also investigated for glass-based actuators pro-

duced using the microfabrication procedure outlined in Section 3.3.2. Two di�erent elec-

trode materials were considered: sputter deposited copper (♦) and sputter deposited

tungsten (O). Tungsten is the metal of choice for many plasma exposed applications due

to its superior thermo-mechanical properties. Tungsten has the lowest sputtering yields

of all metals and low thermal expansion [43], making it an ideal candidate for glass-based

plasma actuators. Due to its resistance to sputtering erosion, tungsten was selected

for comparison with a common electrode material, copper. These cases are compared

against the conventionally constructed 0.18mm thick Kapton actuator (�) from Hanson

et al. [18]. Despite signi�cantly di�erent dielectric thicknesses, the actuators described in

this section were of comparable dielectric constant to thickness ratios.

The time evolution of power consumption, P , e�ective capacitances from the positive

and negative half cycles, C+eff and C−

eff , respectively, as well as cold capacitance, Co is

shown in Figure 6.5 for various actuators. The standard deviation of the measured power

per cycle was typically ≤3%, for the copper on glass actuator while the tungsten actuator

exhibited an average standard deviation of approximately 9%.

The microfabricated actuators with copper and tungsten electrodes exhibited an 8%

decrease and a 7% increase in power consumption, respectively, over a 60 hour actuation

period at 6 kV, 4 kHz. It should be noted that the �nal power consumption per length

values recorded for both the microfabricated actuators were the same, and were within

experimental uncertainty of their respective initial values. This is a signi�cant improve-

ment from a Kapton device which exhibited a 40% increase over a 44 hour period.

As discussed in Section 5.2, the e�ective capacitance extracted from the positive-going

half cycle of the applied signal is substantially di�erent from the e�ective capacitance

extracted from the negative-going half cycle for microfabricated actuators with tungsten

electrodes. This is demonstrated by comparison of tungsten results plotted in Figure 6.5b

and Figure 6.5c. According to Figure 6.5b, C+eff incurs a �uctuating and large increase

(last recorded values is 53% above the initial value), while Figure 6.5c demonstrates a

minor increase in C−eff of only 5% over the entire 60 hour duration. Recall that C−

eff

matched the Ceff values which occur most frequently as determined via the histogram

method. The long-term actuation results demonstrate that for tungsten on glass ac-


P/P0

Ceff/Ceff,0

Co/Co,0

Ceff/Ceff,0

time [hrs]

0 10 20 30 40 50 60

1.4

1.2

1

0.8

1.8

1.6

1.4

1.2

1

0.8

1.1

1.05

1.0

0.95

(a)

(b)

(c)

(d)

+

_

+

_

1.8

1.6

1.4

1.2

1

0.8

KaptonCu on glassW on glass

Figure 6.5: Variations in the (a) power consumption, (b) e�ective capacitance (positivehalf cycle),(c) e�ective capacitance (negative half cycle), and (d) cold capacitance ofvarious actuators during 60 hours of continuous operation at 6 kV, 4 kHz. Values shownare normalized by respective initial measurements at t=0hrs. No data was recordedbetween t=32hrs and t=45hrs for the microfabricated actuator with tungsten electrodes.

tuators, the change in predominant e�ective capacitance is minor, however, during the

positive-going half cycle, the discharge features signi�cant current spikes which a�ect the

determination of the e�ective capacitance values. This e�ect can also be seen in the varia-

tions in power consumption over time as well as in the relatively large standard deviation

in power per cycle. The cold capacitance of the tungsten actuator also demonstrates

variance due to the �uctuations in charge transfer which create greater uncertainty in

the averaged cycles and Q − V cyclogram shapes. Overall, tungsten results indicate a


minor increase in charge transfer over the duration of 60 hours compared to handmade

devices with polymer dielectrics.

Alternatively, copper electroded actuators demonstrate minor and steady decreases

in C+eff and C−

eff of 4% and 2% respectively. The cold capacitance values of the copper

on glass actuator are constant within uncertainty over the test period, although a mild

decreasing trend is observed. The minor decreases in electrical quantities may indicate a

subtle weakening of charge transfer with time.

0 hr 1 hr 4 hr 24 hr 48 hr 60 hr

dielectric3 m

m

0 hr 1 hr 4 hr 24 hr 48 hr 60 hr

dielectric3 m

m

(a)

(b)

exposedelectrode

exposedelectrode

domain

Figure 6.6: CCD images of exposed electrode and dielectric surface for (a)microfabricatedcopper on 0.3mm glass, and (b) microfabricated tungsten on 0.3mm glass actuatorsduring continuous operation at 6 kV, 4 kHz.

Figure 6.6 shows the CCD images of the microfabricated actuator surfaces throughout

continuous operation for 60 hours. The tungsten electrodes in Figure 6.6b showed the

appearance of irregular surface characteristics at the plasma forming edge but relatively

minor material degradation. Copper electrodes, however showed erosion and oxidation in

the presence of plasma over time, which could result in loss of charge transfer capability.

Ozone production is a byproduct of the ionization of air by plasma actuators. Since ozone

is a powerful oxidant, it is capable of oxidizing the metallic electrodes. Copper readily

forms oxides in air alone, thus in the plasma environment, this oxidation is magni�ed.

These oxides are a hindrance to the conductivity of the electrode and add to irregular

surface morphologies. In contrast, tungsten has superior resistance to corrosive agents

and exhibits less surface damage.


0.05 mm

0.05 mm

(a)

(b)

(c)

exposed electrode

plasma-forming edge

0.05 μm

Figure 6.7: Comparison of unused (left) andused (right) plasma actuators via SEM mag-ni�cation for (a) hand-laid Kapton and coppertape, (b) sputter deposited copper electrodeson glass, and (c) sputter deposited tungstenelectrodes on glass.

Handmade 0.18mm Kapton and mi-

crofabricated 0.3mm glass devices were

further compared following 5 hours of

continuous operation at 5 kV, 4 kHz. The

SEM images shown in Figure 6.7 illus-

trate the material quality for both new

and used devices. The exposed electrode

and the dielectric surfaces are shown in

the upper and lower halves of each im-

age, respectively. The SEM images of

the unused Kapton and copper tape ac-

tuator clearly illustrate the imperfections

of hand-cut copper tape and the sig-

ni�cant variations in material surfaces.

Following usage, the degradation of the

electrode, dielectric, and copper tape

adhesive in the plasma-forming region

are visible. The microfabricated copper

and glass actuator also exhibits appre-

ciable degradation of the exposed elec-

trode along the plasma-forming edge af-

ter usage. In comparison, the tungsten

electrode shows limited degradation af-

ter usage. Glass-based devices show no

obvious signs of dielectric degradation at

this magni�cation.

Additional EDS elemental analysis on the used devices identi�ed several points of

interest. Firstly, signi�cant oxygen build-up occurs in all cases at the plasma forming

edge and the downstream plasma region, consistent with exposure to ionized oxygen

during operation. Secondly, trace amounts of metal for both copper and tungsten mi-

crofabricated actuators are found in the downstream plasma region, consistent with a

weak sputtering phenomenon. This is not the case for the handmade actuator, probably

since the foil of the electrode is farther away from the dielectric due to the thickness of

the adhesive, which would make metal deposition less likely. Finally, consequent traces

of copper occur along the plasma forming edge of the microfabricated copper actuator

whereas no trace of tungsten is found on the microfabricated tungsten actuator at the


2 µm 2 µm(a) (b)

Figure 6.8: Comparison of the interface between the plasma-forming electrode edge (top)and the dielectric surface (bottom) for used microfabricated actuators with (a) sputterdeposited copper electrodes and (b) sputter deposited tungsten electrodes via SEM mag-ni�cation.

plasma forming edge.

Figure 6.8a and Figure 6.8b further demonstrate the signi�cant di�erences in surface

morphologies at the plasma-forming interface for copper and tungsten actuators, respec-

tively, following 5 hours of operation at 5 kV, 4 kHz. The tungsten electrode surface

upstream of the interface appears homogeneous, even at high magni�cation, whereas the

copper electrode su�ers from surface irregularities and visible deposits. At the electrode-

dielectric interface, the oxidation of the tungsten electrode features a more compact

structure compared to that of copper oxidation. Metals such as calcium, magnesium,

tungsten, and uranium form porous oxides which oxidize at a linear rate [2, 28, 53]. In

contrast, metals such as cobalt and copper form non-porous oxides and obey parabolic

laws of oxidation [53]. This accounts for the relatively severe oxidation of copper elec-

trodes. Furthermore, the accumulation and deposition of oxides on the copper electrodes

e�ects a much larger surface area (encroaching onto the dielectric) than for tungsten

electrodes. The important di�erence between these two electrode materials is the com-

parative e�ectiveness with which tungsten retains its surface properties.

The di�erence in how the electrode materials degrade a�ects the quality of the plasma

generation, as demonstrated in Figure 6.9 for both copper and tungsten electrodes. The

�gure shows images of the plasma generated at di�erent excitation voltages produced

by modulating a 5 kV, 4 kHz waveform with another sine wave of amplitude 1 and 25 s


period. Both sets of actuators were operated at 5 kV, 4 kHz prior to visual recording. The

tungsten electrode shown has uniform plasma formation of gradually increasing strength

in response to increasing applied voltage. In contrast, the copper electrode produces a

more �lamentary discharge, as noted by bright and dim spots in the plasma. Plasma

generation by the copper electrode also occurs abruptly, as seen between the third and

fourth frames (Figure 6.9a). Predictable and uniform discharge behaviour is preferable

from a �ow control standpoint as well as in terms of actuator robustness. Filamentary

discharge can lead to irregularities in induced �ow and localized stresses on the device,

which in turn can lead to premature failure.

a)

b)

4.03 4.23 4.41 4.56 4.80 5.00

Figure 6.9: Plasma generation response to increasing applied voltage by a used microfab-ricated actuator with (a) sputter deposited copper electrodes and (b) sputter depositedtungsten electrodes. The applied voltage amplitude is listed above each image column inkV.

7 | Electromagnetic Radiation

from Plasma Actuators

Despite extensive in-laboratory usage, previous plasma actuators studies have neglected

to evaluate the severity and extent of electromagnetic radiation generated by plasma

actuator operation. This information is relevant to protect those in environments where

plasma actuators may be implemented for �ow control and, more immediately, for the

safety of plasma actuator researchers in laboratory environments, as well as their equip-

ment. The electric and magnetic �elds surrounding a plasma actuator were documented

for a variety of operating conditions as part of the present work. The values shown here

are the resultant of RMS measurements taken along three perpendicular axes as de�ned

in Section 4.4 using a ME3951A �eld meter. The measurements were taken at distances

from the actuator centre for various orientations as shown in Figure 4.5. The uncertainties

in resultant electric and magnetic �eld strengths were calculated in quadrature according

to [61] using individual component error of ± 2% as per the �eld meter speci�cations for

both sensors. Uncertainty in the radial distance to the actuator centre was estimated

at ± 4%. The �eld strengths and associated uncertainties were used to determine the

curve �ts shown in the following sections using a weighted least squares regression. The

background electric and magnetic �eld strengths in experimental area were 0V/m and

0 nT, respectively.

7.1 |Electric Field

As shown in Figure 7.1, the electric �eld strength contours are approximately the same

for all orientations examined, indicating that the strength is largely symmetric radially

about an operating plasma actuator.

Using the approximation that the �eld about the actuator is independent of ori-

entation, the mean electric �eld for each position was calculated over all orientations.

46

Chapter 7. Electromagnetic Radiation from Plasma Actuators 47

0

100

200

300

400

500

600

700

800

900electric fi

eld stren

gth [V

/m]

hei

ght

[m]

4 kV

8 kV

10 kV

12 kV

1.5

1

0.5

00.5 1 1.5 2

distance [m]

0 45 90315 270o oo o o

1.5

1

0.5

0

1.5

1

0.5

0

1.5

1

0.5

0

0.5 1 1.5 2 0.5 1 1.5 2 0.5 1 1.5 2 0.5 1 1.5 2no data

Figure 7.1: The resultant electric �eld strength at all orientations in response to variousoperating voltages (at 4 kHz).

0 50 100 150 200 250 300

position average electric field [V/m]

dev

iation

[V

/m]

90 o

45o

0o

315o

270o

50

40

30

20

10

0

-10

-20

-30

-40

-50

Figure 7.2: The deviation of each measure-ment from the position averaged electric �eldstrength for all operating conditions.

The deviations of individuals measure-

ment from the location averaged elec-

tric �eld strength are shown graphically

in Figure 7.2 for an operating voltage of

12 kV, 4 kHz. On average, measurements

taken at the 90◦ orientation were 10%

higher than the location averaged �eld

strength. This tendency can be intuited

by considering the location of the high-

voltage electrode. In the 90◦ orientation,

the distance from the �eld meter to the

source is smallest and free of obstructions.

The average deviations of all other orien-

tations were ≤ 7%.

From the contours plots in Figure 7.1 two major conclusions can be drawn. Firstly, the

electric �eld strength, | ~E|, is a function of the radial distance, r, from the actuator centre

regardless of position about the actuator. Secondly, the increase in operating voltage

increases the strength of the resultant electric �eld. These results are more explicitly

expressed in Figure 7.3. For various operating voltages, the trend | ~E| ∝ 1r2

was found,


consistently over all orientations tested. This is the same result one would expect with

the electric �eld generated by a charged point source. Since the electric �eld strength

is dependent on the radial distance from the actuator, the e�ect of various operating

voltages was compared at speci�c grid coordinates, the radial distances of which are

displayed in Figure 7.3b. The linear relation | ~E| ∝ Vpp was found, consistently, for all

orientations tested. The e�ect of operating frequency on resultant �eld strength was also

investigated. No signi�cant correlation was observed between electric �eld strength and

operating frequency.

0.5 1 1.5 2 2.5

radial distance, r [m]

elec

tric

fiel

d s

tren

gth [V

/m]

0

200

400

600

800

1000

r , 4 kV−2.3 0.1

r , 8 kV−2.1 0.2

r , 10 kV−2.0 0.3

r , 12 kV−2.1 0.3

4 6 8 10 12

voltage, Vpp [kV]

elec

tric

fiel

d s

tren

gth [V

/m]

0

100

200

300

400

500

600

V , 0.7 mpp

0.9 0.8

V , 1.0 mpp

0.7 0.4

V , 1.4 mpp

1.1 0.3

V , 1.6 mpp

0.9 0.2

V , 1.8 mpp

1.0 0.2

(a) (b)

Figure 7.3: The resultant electric �eld strength at 90◦ orientation as a function of (a)radialdistance from actuator centre for various operating conditions and (b)operating voltageat various radial locations (at 4 kHz).

7.2 |Magnetic Field

From the contour plots of the resultant magnetic �eld strength, | ~B|, shown in Figure 7.4,

it is shown that the magnetic �eld induced by the plasma actuator is quite weak, dropping

to essentially zero by a radial distance of approximately 1.5m. Figure 7.5 shows magnetic

�eld strength as a function of both radial distance from the actuator and operating

voltage. The trends | ~B| ∝∼1r2

and | ~B| ∝ Vpp were found via weighted least squares

regression, consistently over all orientations tested. These relations mimic those of the

electric �eld strength, as expected as a result of the relationship between electric and

magnetic �elds, | ~B| ∝ | ~E|.It should be noted that operation of the plasma actuator at 4 kV failed to produce

visible plasma discharge. This falls under the heading of `dark discharges' where part


0

5

10

15

20

25

30m

agnetic fi

eld stren

gth [n

T]

hei

ght

[m]

distance [m]

4 kV

8 kV

10 kV

12 kV

0 45 90315 270

1

0.5

0

1

0.5

0

1

0.5

0

1

0.5

0 0.5 1 0.5 1 0.5 1 0.5 1 0.5 1

o oo o o

Figure 7.4: The resultant magnetic �eld strength at all orientations in response to variousoperating voltages (at 4 kHz).

mag

net

ic fi

eld s

tren

gth [nT

]

mag

net

ic fi

eld s

tren

gth [nT

]

0.5 1 1.5 2 2.5

radial distance, r [m]

0

2

4

6

8

10

12

r , 4 kV−2.3 0.2

r , 8 kV−2.2 0.2

r , 10 kV−2.4 0.2

r , 12 kV−2.4 0.2

0

5

10

15

20

25

30

35

40

V , 0.2 mpp

1.0 0.1

V , 0.5 mpp

1.0 0.1

V , 0.54 mpp

0.9 0.1

V , 0.7 mpp

1.1 0.1

4 6 8 10 12

voltage, Vpp [kV]

(a) (b)

Figure 7.5: The resultant magnetic �eld strength at 90◦ orientation as a function of(a)radial distance from actuator centre for various operating conditions and (b)operatingvoltage at various radial locations (at 4 kHz).

ionization of the gas and electron avalanche occur for operating voltages below the break-

down voltage of the gas. From both the electric �eld and magnetic �eld contour sets, as

well as Figure 7.3 and Figure 7.5, it can be seen that operation of the plasma actuator

in the dark discharge regime does not di�er from operation in the glow discharge regime


with respect to electric and magnetic �eld behaviour. This emphasizes the importance of

EMF awareness for researchers working with or in close proximity to plasma actuators, as

even operation below breakdown �eld strength can generate measurable EMF radiation.

7.3 |Concerns with Electromagnetic

Radiation Exposure

The increased usage of technological communicative devices has led to growing concerns

over the potential health e�ects caused by exposure to electromagnetic radiation. Of

particular interest are the health e�ects of exposure to radio-frequency (RF) sources.

Radio-frequency sources operate at frequencies between 3 kHz and 300GHz on the elec-

tromagnetic spectrum. This frequency range supports many widely used applications

including but not limited to radar, satellite, navigation, wireless, cellular, and television

devices [21]. There is great di�culty in quantifying human health damage caused by

EM radiation due to the challenges of isolating sources, as well as the fact that damage

caused by EMF radiation, if any, is non-acute in nature. Studies on the potential health

e�ects associated with RF radiation have examined concerns such as DNA damage, tu-

mour promotion, human cancers, behaviour and cognitive functions, gene and protein

expression, immune response, and reproductive functions [17]. Exposure to RF �elds is

known to induce internal body currents and energy absorption in tissues [21]. Despite

numerous studies on a large variety of health e�ects of RF energy, adverse e�ects are

predominantly related to the occurrence of tissue heating for frequencies from 100 kHz to

3GHz and excitable tissue stimulation from acute exposures for frequencies from 3 kHz

to 100 kHz [17].

Health Canada's Safety Code 6 [17] provides outlines for the limits of human exposure

to RF electromagnetic energy in the frequency range from 3 kHz to 300GHz, based on

continuous review of published scienti�c studies. According to the Safety Code 6, there

exists no scienti�c evidence of chronic and/or cumulative health risks from RF energy

at levels below speci�ed limits provided by the code. Safety Code 6 de�nes controlled

environments as those in which RF �eld intensities have been characterized by means

of measurement, calculation, or modeling with appropriate software and exposure is

incurred by persons aware of potential for, intensity of, health risks associated with, and

mitigation strategies for RF exposure in their environment. Any situations that do not

meet these criteria are considered uncontrolled environments in which RF energy has

been insu�ciently assessed and/or where persons within environment have not received


Table 7.1: Field strength limits set by Health Canada's Safety Code 6.

Environment Type | ~E| [V/m] | ~B| [nT]Controlled (3 kHz - 1000 kHz) 600 6159Uncontrolled (3 kHz - 1000 kHz) 280 2752

adequate RF awareness training and lack the means to asses or mitigate their exposure

to RF energy. The maximum electric and magnetic �eld strengths for both controlled

and uncontrolled environments are listed in Table 7.1.

For the highest operating voltage tested, 12 kV, the controlled environment electric

�eld limit of 600V/m was exceeded between 0.5m and 0.7m of the actuator, while the

uncontrolled environment limit of 280V/m was exceeded within approximately 1.0m of

the actuator. For the lowest operating voltages tested, 4 kV, the controlled environment

electric �eld limit of 600V/m was exceeded with between 0.2m and 0.5m of the actuator,

while the uncontrolled environment limit of 280V/m was exceeded within approximately

0.5m of the actuator. The largest resultant magnetic �eld recorded was 0.6% and 1.4%

of the controlled and uncontrolled environment magnetic �eld limits, respectively.

For the actuator used in EMF tests, maintaining a distance of 1.5m guaranteed EMF

radiation exposure below the limits set forth by Health Canada, for the range of operat-

ing conditions considered here. A simple EMF survey was recorded at the location of the

author's workstation chair for the PIV experimental set-up. In this set-up, operation of

a microfabricated actuator at 10 kV, 4 kHz generated an electric �eld of approximately

44V/m at a radial distance of approximately 2.55m from actuator centre. This location

represented the approximate location of the head of a person seated at the PIV com-

puter. The electric �eld strength was found to decrease at vertical positions closer to

the �oor since the actuator was approximately at head level. Reassuringly, this reading

was signi�cantly below the recommended limits for exposure to EMF radiation stated in

Safety Code 6. Although, the workspace is considered a safe distance from the operating

actuator according to Health Canada, the ME3951A user manual recommends exposure

limits of 1V/m (electric �eld) and 20 nT (magnetic �eld) for frequencies above 2 kHz in

areas where people spend substantial amounts of time, such as the workplace. As such,

for the experiments presented in the current work, prolonged exposure to any operating

plasma actuators was minimized and ample distance was kept between persons and live

actuators. Evaluation of EMF exposure is an important consideration in experimental

set-up design for health and safety of both humans and equipment.

Aside from human safety concerns, EMF radiation can also pose a risk to damaging

electronic lab equipment through static build-up. AC electric �elds can induce currents in


conductive materials in the �eld. Accumulation of potential on a conductive surface, such

as an ungrounded metallic equipment casing, can result in electrostatic discharge (ESD)

with grounded (or lower potential) objects, such as a human body. ESD can cause damage

to delicate electronic components. For this reason it is important that all lab equipment

be properly grounded in the vicinity of an operating plasma actuator. EMF radiation

can also in�uence sensitive measurement tools to give corrupt readings, for example

thermocouples or the strain gauges of an electronic balance. Care should be taken when

plasma actuator experiments occur in the same workspace as other experiments for the

sake of both lab equipment and the safety of other researchers who may not be aware of

the EMF radiation safety standards.

8 | Summary & Conclusions

In the present study, two methods for the mitigation of plasma-induced degradation as

described in Section 3.3 were proposed and assessed. These methods were a protective

surface treatment suited for degradation prone polymer-based actuators and a micro-

fabrication technique for precise construction on plasma-resistant glass dielectric. The

electrical quantities of the constructed devices were determined using a probe capacitor

via charge-voltage cyclograms (Lissajous �gures). Microfabricated actuators were char-

acterized by their electrical properties, as well as by velocity performance as determined

using PIV. Actuators employing the proposed methods for increased robustness were

compared with a conventionally constructed plasma actuator over extended periods of

operation. The degradation of these devices was assessed by changes in electrical char-

acteristics, visual analysis, SEM images, EDS analysis, and for microfabricated devices,

plasma generation. Two di�erent electrode materials were explored for microfabricated

actuators, copper and tungsten. The electric and magnetic �elds surrounding an operat-

ing plasma actuator were characterized and the health and safety implications discussed.

The main �ndings are summarized in the following sections.

8.1 |Investigation of Microfabricated Devices

In agreement with the work of Hanson et al. [18], the power law P ∝ f 1.2 was found

between power consumption and frequency of the applied signal for microfabricated ac-

tuators. Within uncertainty of the often reported power law relation, P ∝ V 3.5pp , the power

laws for microfabricated devices with copper and tungsten electrodes were P ∝ V 3.8pp , and

P ∝ V 3.3pp , respectively. The maximum induced jet velocities were found to follow the

relations umax ∝ P 0.74, umax ∝ f 0.8, and umax ∝ V 2.5pp for actuators with copper elec-

trodes and umax ∝ P 0.75, umax ∝ f 0.7, and umax ∝ V 2.7pp for actuators with tungsten

electrodes. The power laws found for the umax-voltage relation are in agreement with

those reported by other studies in the literature. These �ndings verify that the microfab-

ricated devices operate according to established relations for conventionally constructed

53

Chapter 8. Summary & Conclusions 54

(handmade) plasma actuators, and therefore direct comparison of electrical properties is

justi�ed. However, one key di�erence was noted in the current traces recorded for glass-

based actuators. The current response of glass-based microfabricated actuators to the

positive-going half cycle of the applied signal featured large spikes which indicate sudden

incidents of elevated charge transfer. This generated altered cyclogram shapes compared

to typical Kapton cyclograms and increased the uncertainty and variance in electrical

quantities determined from the plots. This e�ect was more noticeable for actuators with

tungsten electrodes. The velocity pro�les of microfabricated actuators, a PDMS treated

actuator, and a typical untreated Kapton actuator were compared. All devices produced

laminar wall jets.

8.2 |Degradation of Plasma Actuators

Both the application of PDMS surface treatment for Kapton actuators and the micro-

fabrication of glass actuators were found to improve the robustness of plasma actuators

compared to a conventional untreated Kapton actuator. Time evolutions of power con-

sumption, e�ective capacitance, and cold capacitance over 60 hours of continuous op-

eration were correlated to visual analysis of the physical degradation of the actuators.

PDMS treatment of Kapton actuators was found to alter the typical degradation patterns

associated with polyimide erosion. Rather than uniform dielectric degradation along the

plasma-forming edge, the degradation occurred at weak spots in the Kapton surface and

developed into sporadically occurring coral-like degradation patterns. The onset of these

patterns coincided with increases in power consumption and e�ective capacitance. The

presence of the PDMS treatment delayed and reduced the rate of change in these elec-

trical properties compared to the untreated Kapton actuator. Additionally, it was found

that repeated PDMS application could prevent the occurrence of polyimide degradation

and prolong the actuator lifetime.

Alternatively, microfabricated actuators o�er a greater degree of geometric precision

and were also found to have superior resistance to plasma-induced degradation com-

pared to handmade counterparts. Sputter deposited tungsten electrodes were found to

exhibit superior degradation resistance and plasma generation following usage compared

to sputter deposited copper electrodes. SEM and EDS analysis demonstrated that copper

electrodes undergo erosion, sputtering, and oxidation to a greater extent than tungsten

electrodes. However, tungsten electrodes provided more variance in electrical quantities

than copper electrodes due to signi�cant spikes in the current response. The glass di-

electric exhibited no signs of degradation. Microfabrication was found to be an e�ective


means to enhance repeatability and actuator robustness.

8.3 |Electromagnetic Field Considerations

The electric and magnetic �elds surrounding an operating plasma actuator were charac-

terized at various voltages and frequencies. The resultant electric �eld strength generated

by a plasma actuator was found to be a function of radial distance, described by | ~E| ∝ 1r2,

the same relation which applies to a point charge. The electric �eld was directly pro-

portional to the applied voltage, | ~E| ∝ Vpp. For magnetic �elds, the relations | ~B| ∝∼1r2

and | ~B| ∝ Vpp were found, as to be expected from the relationship between electric and

magnetic �elds (| ~B| ∝ | ~E|). No correlation was found between the applied frequency and

the electric or magnetic �elds. Interestingly, it was found that the electromagnetic �eld

behaviour was unchanged for a plasma actuator operating in the `dark discharge' regime

(no visible plasma), emphasizing the importance of EMF analysis in the lab space. Even

for operation below breakdown �eld strength, EMF strength exceeded the exposure lim-

its found in Safety Code 6 developed by Health Canada. These exposure limits, human

health e�ects, and lab equipment concerns are discussed in Section 7.3.

8.4 |Concluding Remarks

The present work aims to enhance plasma actuator manufacturing techniques for more

precise and robust devices. Two avenues for degradation reduction were explored with

success. The degradation mitigation methods proposed here have di�erent applications.

For researchers that are more conceptually focused, using a protective treatment to pro-

long the usable lifetime of conventional handmade actuators, makes perfect sense. The

combined variability of handmade devices and oil application make for inconsistency and

imprecision amongst actuators. Thus, this method for increased plasma-resistance is best

suited for experimental use where consistency amongst multiple actuators is not neces-

sary and where simple and inexpensive actuator construction is appropriate, such as a

proof-of-concept style investigations. Alternatively, for the future application and device

development focused researchers, microfabrication elevates the quality of the manufac-

tured devices and will lead to further advancements in plasma actuator manufacturing.

This method of construction is better suited to applications which require multiple iden-

tical actuators, are subject to prolonged continuous operation, have poor/limited access

to the actuator (such that reapplication of a surface treatment is an inviable option),

and/or where the expense is justi�ed. Examples include �ow control experiments involv-


ing multiple plasma actuators of speci�c functionality where precision and reproducibility

are paramount.

In the future, it is recommended that further performance characterization of glass-

based microfabricated devices be explored, including the changes in velocity �eld mea-

surements at various times during prolonged continuous operation. Monitoring the degra-

dation in momentum transfer should provide further insight to the e�ectiveness of these

microfabricated actuators. Other potential avenues to explore include the optimization

of electrode thickness with respect to robustness and jet output, quanti�cation of the

consistency in discharge speci�c quantities (P , Ceff , Co) between microfabricated actu-

ators, and the e�ects of increasing electrode edge irregularities at various scales. The

latter suggestion aims at investigating di�erences in actuator function, performance,

and/or degradation patterns due to reduced crispness or straightness of the electrode

edge. From this study, it could be determined if photolithography is required to achieve

maintenance of functionality, or if a physical mask (stencil) could be used to produce

actuators of comparable quality (for reduced time and �nancial investment).

Glass (εglass = 6.2) has higher dielectric constant than Kapton (εKapton = 3.9) and

other commonly used dielectrics. Therefore, at the same thickness, glass-based actuators

have a higher e�ective capacitance, which raises the localized concentration of electric

�eld lines. This has the equivalent e�ect of increasing the current density which pro-

motes the formation of streamers or �lamentary discharge at lower voltages [62]. This is

a potential contributer to the spikes in charge transfer observed in glass-based devices.

Glass is also delicate and di�cult to handle at the �ne thicknesses used in the present

work. Ideally, the microfabrication technique can be employed using a di�erent dielec-

tric substrate that has a lower dielectric constant, improved handling durability, and

adequate plasma resistance. The procedure could be adapted using E-beam evaporation

deposition, a lower temperature process relative to sputtering at deposition thicknesses

below a micron, should temperature sensitivity become an issue with future dielectrics.

Furthermore, the microfabrication process should be adapted for the deposition of gold

electrodes, as gold is resistant to oxide-�lm formation at temperatures below its melting

point (1063◦C) [59], unlike the metals used in the current work which both oxidize in air

at room temperature [53, 54].

In conclusion, it is hoped that the work presented here contributes to the advance-

ment of plasma actuator manufacturing and o�ers solutions to commonly occurring, but

rarely addressed degradation issues. With this work, the author aims to emphasize the

signi�cance of material degradation and the impact it has on actuator function, as well

as to encourage the ongoing quest for suitable actuator materials within the community.

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