III. Thermomechanical Analysis of SU8 Thermal Actuator

Fabrication and Non-linear Thermomechanical
Analysis of SU8 Thermal Actuator
Leema Rose Viannie, G. R. Jayanth, V. Radhakrishna and K. Rajanna
 Abstract—SU8 based micromechanical structures are widely
used as thermal actuators in the development of compliant
micromanipulation tools. This paper reports the design, nonlinear thermomechanical analysis, fabrication, and thermal
actuation of SU8 actuators. The thermomechanical analysis of
the actuator incorporates non-linear temperature-dependent
properties of SU8 polymer to accurately model its thermal
response during actuation. The designed SU8 thermal actuators
are fabricated using surface micromachining techniques and
electrical interconnects are made to it using flip-chip bonding.
The issues due to thermal stress during fabrication are discussed
and a novel strategy is proposed to release the thermal stress in
the fabricated actuators. Subsequent characterization of the
actuator using a 3D optical profilometer reveals excellent
thermal response, good repeatability and low hysteresis. The
maximum deflection is about 10 μm for an actuation current of
about 5 mA. The experimentally obtained deflection profile and
the tip deflection at different currents are both shown to be in
good agreement with the predictions of the non-linear
thermomechanical model. This underscores the need to consider
nonlinearities when modeling the response of SU8 actuators.
nonlinearthermomechanical analysis, 3D optical profiler, residual
MEMS fabrication technology has enabled dramatic
miniaturization of mechanical sensors and actuators [1], [2].
Micromanipulators developed from these miniature devices
can be used to precisely hold and assemble micro-parts such
as micro-gears [3] and micro-mirrors [4]. In biological
applications, micromanipulators are often used to capture,
immobilize and characterize individual living cells [5], [6].
Such micromanipulators or micro robotic arms are often
integrated with an actuator in order to produce controlled
mechanical motion. These actuators can be driven using
electrostatic [7], piezoelectric [8], shape memory [9] or
thermal [10] actuation techniques. Among these, thermal
actuation has the advantages of large actuation force, low
driving voltage and large mechanical deflection [11]. The
performance of thermal actuators mainly depends on the
coefficient of thermal expansion (CTE) of its constituent
materials. Most often, MEMS based thermal actuators are
fabricated from silicon [12]. Being relatively stiff and
possessing small CTE, silicon produces small mechanical
Leema Rose Viannie, G. R. Jayanth and K. Rajanna are with the Department
of Instrumentation and Applied Physics, Indian Institute of Science,
[email protected];
[email protected]; [email protected]).
V. Radhakrishna is with Space Astronomy Group, ISRO Satellite Centre,
Bangalore, India. (e-mail: [email protected])
deflections. The development of micromechanical tools for
single cell manipulation and characterization; positioning and
moving of cells or parallel manipulation of large number of
biological entities require that the actuators must be
biocompatible and capable of generating large mechanical
displacement for small driving voltages at lower operating
temperatures [6]. Therefore, research groups have
investigated fabrication of thermal actuators by alternate
materials such as polymers, since they possess low Young’s
modulus and high CTE. Some commonly used polymers in
the fabrication of thermal actuators are parylene C [13],
polyimide [14], polypyrrole [15] and SU8 [16]. Among them,
SU8 polymer is particularly suited for fabricating high
aspect-ratio MEMS devices [17]. It is an epoxy based
negative photoresist which can be easily patterned using
conventional UV photolithography techniques. Fully crosslinked SU8 polymers have several desirable properties
including low Young’s modulus, better chemical resistance
and excellent biocompatibility [18]. Due to its high CTE
value, SU8 thermal actuators provide large mechanical
deflections at lower operating temperatures [6], [19]. While
the low Young’s modulus reduces the stiffness and hence the
bandwidth of actuation, SU8 polymer actuators find
applications in several areas which do not require large
bandwidth and but do require lower actuation force [5], [6],
[19], such as in the development of micro-valves [8], microextractors [20], micro-actuators [21], and micro-cages [22].
These devices find applications in areas such as
characterization of biological cells and tissues. Likewise, they
can also be used as probes in Atomic Force Microscopy,
wherein they can be employed for performing quasi-static
force spectroscopy and indentation experiments [23],
especially on soft, biological samples such as macromolecules and cells.
In order to evaluate the performance of SU8 thermal
actuators at elevated operating temperatures, a precise
understanding of thermomechanical properties of SU8 is
essential. Earlier reports have investigated the mechanical
[24] and thermal [25] properties of fully cross-linked SU8
polymer for MEMS applications. These properties are used in
several studies that adopt analytical and finite element (FE)
modeling of SU8 thermal actuators [19], [25], [26], [27],
[28]. The FE modeling assumes linear thermo-elastic
behavior of the polymer, and therefore neglects the variation
of Young’s modulus and CTE at all operating temperatures.
These models are used to estimate the temperature
distribution in a heated thermal actuator. However, a recent
report by S. Chung et al [29] revealed the effects of
temperature on the mechanical properties of SU8 polymer. It
was found that the Young’s modulus of SU8 shows strong
non-linear temperature dependence. Therefore, in this work
we incorporate the effect of temperature on the Young’s
Modulus, evaluate the actuator’s thermomechanical response
and verify the same experimentally.
In this paper, we present the design, nonlinear
thermomechanical analysis, fabrication and actuation of a
polymeric thermal actuator. The actuator consists of SU8
microcantilever with an integrated Au thin film resistor.
Nonlinear finite element analysis (FEA) has been used to
simulate the thermomechanical response of the actuator by
incorporating the dependence of Young’s modulus and
Poisson’s ratio on temperature. The SU8 thermal actuator is
fabricated using surface micromachining techniques.
Subsequently, the thermal actuation of the actuator has been
experimentally characterized using 3D optical profiler. A
novel strategy is proposed to release the thermal stress in a
fabricated actuator. Subsequent experiments show good
agreement between the measured deflections of the actuator
and that of the nonlinear FEA. This underscores the need to
consider the effect of temperature on the performance of SU8
based thermal actuators.
The rest of the paper is divided as follows: section II
discusses the design of SU8 thermal actuator. The analytical
model for steady state temperature distribution, finite element
analysis (FEA) of a thermally heated actuator along with the
non-linear thermomechanical analysis of the SU8 actuator is
discussed in section III. A detailed account of the actuator
fabrication and its electrical interconnection is discussed in
section IV. This section also discusses the challenges faced
during the fabrication process. The experimental set-up used
to measure the motion of the actuator is mentioned in section
V. Finally, a comprehensive discussion of the experimental
results and their comparison with finite element analysis is
presented in section VI.
When current is passed through the resistor, its temperature
rises due to Joule heating. This rise in temperature induces
internal thermal stresses within SU8-Au bimorph. Since the
thermal expansion of SU8 (α2) is greater than Au (α1) it
produces a bending moment in SU8-Au bimorph thereby
resulting in the mechanical deflection of SU8 cantilever.
Also, the direction of cantilever deflection is towards lower
CTE layer i.e. towards the Au layer. The maximum tip
deflection of the cantilever bimorph is related to the
temperature distribution within the cantilever, its physical
dimensions as well as its material properties [30]. Since the
temperature varies along the length of the cantilever, and also
the Young’s modulus of SU8 polymer changes with
temperature [29], the deformation profile for such a nonlinear thermal actuation phenomenon cannot be estimated
accurately using simple analytical models. Therefore, we
employ finite element analysis to evaluate the
thermomechanical response of SU8 thermal actuator.
Young’s Modulus
Thermal conductivity
Young’s Modulus
Thermal conductivity
Fig. 1(a) and (b) shows the geometry of the SU8 thermal
actuator comprising of a rectangular SU8 microcantilever
with an integrated gold (Au) thin film resistor. The actuator’s
dimensions and material properties are given in Table I. In
this design, the two structural layers namely the SU8 polymer
and the Au thin film constitute a thermal bimorph. The Au
resistor, which is symmetrically patterned over the SU8
cantilever, serves as a resistive heating element.
This section discusses the temperature distribution and the
corresponding deformation profile of a SU8 thermal actuator.
Section III A proposes a simple model to obtain the
temperature distribution along the length of the cantilever,
and validates it by means of finite element analysis. Section
III B employs finite element analysis to obtain the
deformation profile of the actuator.
A. Steady-state temperature distribution on the thermal
Fig. 1. (a) Schematic cross-section of SU8 thermal actuator showing SU8-Au
bimorph structure. (b) Schematic showing the top view of rectangular SU8
microcantilever with integrated Au resistor.
In order to determine the steady-state temperature
distribution on the thermal actuator, a one dimensional (1D)
analytical model was formulated. The model assumes the
presence of conductive heat transfer [31] within the gold
layer and between the gold and polymer layers, but ignores
the effects of convective and radiative heat transfer. With
these assumptions, the variation of temperature Tg along the
contour of the gold resistor depends on the actuating current
I and the polymer temperature Tp as,
d 2Tg
 k w 
I 2
  2 2  (Tg  T p )
1 11
 k1w1t1t2 
where, x1 is a distance variable along the contour of the
resistor. The other symbol definitions are mentioned in Table
I. Likewise, by applying the steady-state heat equation, the
variation of polymer temperature Tp along the contour of the
gold resistor can be written as
d 2T p 2
 (Tg  T p )  0
dx12 t22
By assuming that the ends of the Au resistor and the polymer
layer at 1 = 0 and 1 = 1 are maintained at zero, i.e.,
x 0
 Tg
x  L1
 0 and Tp
x 0
 Tp
x  L1
 0 , Eqn. (1) can
be solved analytically to obtain the variation of temperature
along the contour of the Au resistor. This is given by
T ( x1 ) 
 p 1  2
me x1  ne  x1  
 x1   r  x1   m  n  
a 
 4a
 a
 2a
  L1
 L1
  L1
1  e 
 L1
 L1
 L1
2 
The maximum temperature rise within the Au resistor occurs

in the middle of the resistor. Hence Eqn. (3) at 1 = 21 can be
written as
Fig. 2(a) shows the steady-state temperature distribution in
the SU8 actuator obtained from FEA. Fig. 2(b) compares the
temperature distribution in the Au resistor obtained from
finite element analysis (FEA) with that obtained analytically
by using Eqn. (3). It is seen that the maximum difference
between the two is about 0.38%. The close match between
the two reveal that Eqn. (3), obtained from the ID model,
provides simple closed-form expression for temperature
distribution that enables quickly and accurately estimating the
steady-state temperature distribution. Further, the functional
dependence on the parameters evident in Eqns. (3) and (4)
facilitates design and optimization of the actuator. Finally, the
results of FE simulation also reveal that the effect of
convective heat transfer is negligible and thus, it does not
significantly affect the temperature profile.
B. Non-linear thermomechanical analysis of the SU8 thermal
actuator using FEA
me L1  ne L1  (m  n)   p  1 L1 
Fig. 2. (a) FE simulation showing the steady state temperature distribution in
a heated SU8 thermal actuator. (b) Analytical and finite element analysis
(FEA) showing the variation of temperature along the length of the SU8
thermal actuator.
   L1  
q    L1 
   p  1  L1
  me   ne
  r  1   m  n 
  2  4
a 
In order to validate this model, the steady state temperature
distribution within the SU8 thermal actuator was simulated
using finite element software (COMSOL). A 3D model of the
actuator was constructed using the dimensions and material
properties as mentioned in Table I. In this simulation, the
convective heat transfer for air was assumed to be 5 W/m2K.
Also, the effect of radiation was ignored considering low
operating temperatures in polymeric devices.
The thermal actuation of the SU8 thermal actuator was
simulated using coupled field multiphysics in COMSOL, by
combining the effects of Joule heating and thermal expansion
with mechanical deformation. Since the material properties of
a fully cross-linked SU8 polymer change with temperature,
the analysis also incorporated this effect. The dependence of
Young’s modulus and Poisson’s ratio on temperature for SU8
was obtained from S. Chung et al [29], who report their
values at four distinct temperatures, viz., 25ºC, 50ºC, 100ºC
and 150ºC. Since the SU8 material softens at 200˚C and loses
its mechanical stability [17], the Young’s modulus at 200ºC
was assumed to be zero. Subsequently a suitable nonlinear fit
was performed in order to estimate the mechanical properties
at other intermediate temperatures (Fig. 3(a) and (b)).
Accordingly, the dependence of Young’s modulus on
temperature was modeled to be
E (T )  E0 exp  T 
Where the constant E0 was identified to be 5.76GPa and the
temperature constant, Tc was identified to be 55.82˚C.
Likewise, the dependence of Poisson’s ratio on temperature
was modeled as
 (T )   0 1T  2T 2
For SU8, the constants0 , 1 and 2 which are obtained from
the experimental results were found to be 0.3116,
0.00103ºC-1and −1.99 × 10-6ºC-2respectively. The 3D CAD
model of the SU8 cantilever with Au strip over its surface is
shown in Fig. 3(c). The deformation of the actuator obtained
by incorporating the temperature effects on mechanical
properties in the finite element model is seen in Fig. 3(d). The
deflection profile as seen in Fig. 3(c) illustrates the direction
of cantilever bending towards the lower CTE layer i.e.
towards the Au layer along the positive Z-axis. The nonlinear thermomechanical model predicts the deflection profile
of the SU8 thermal actuator for actuation currents. The results
of this finite element analysis (FEA) are discussed further in
Section V(C), Fig. 13(b).
deposition, a bilayer lift-off technique was adopted which
includes lift-off resist (LOR10A) along with a positive
photoresist (S1813). By patterning and developing the bilayer
resist, an undercut is formed. Metal film deposited in such a
manner does not touch the walls of the resist and later the
resist can be easily stripped off. (d) 5µmthickSU8cantilever
layer was defined over the patterned metal resistor and the
contact pads by spinning negative photoresist SU8 2005 at
4000 rpm. Pre-exposure baking was done at 95ºC for 2 min
followed by 105 mJ/cm2of UV exposure. Post-exposure
baking was done at 95ºC for 3 min. (e) 100µm thick SU8
base layer was patterned over SU8 cantilever patterns. The
thick SU8 base serves as a supporting structure that holds
together the cantilever and electrical contact pads and also
helps in safe handling of the suspended SU8 thermal actuator.
In order to obtain 100µm thick base layer, SU8 2035 was
spun over SU8 cantilever patterns at 1000rpm followed by
pre-exposure baking at 65ºC for 5 min and 95ºC for 25min.
UV exposure of 450mJ/cm2results in SU8 base patterns. This
was followed by post exposure baking which was carried out
at 65ºC for 10 min and 95ºC for 20 min. At the end of the
photolithography process, the unexposed SU8 resist was
developed by immersing the entire silicon wafer containing
the SU8 cantilever and the base patterns in SU8 developer
(Microchem) for about 10min.
Fig. 3. (a) Variation of Young’s modulus of SU8 with temperature. (b)
Variation of Poisson’s Ratio of SU8 with temperature. (c) 3D CAD model of
the SU8 thermal actuator indicating the position of the Au resistor. (d) FE
simulation showing thermomechanical deflection of the actuator by
incorporating temperature dependent properties of SU8.
The SU8 thermal actuators were fabricated using surface
micromachining technique. Section IV A discusses the
patterning of SU8 cantilever structures along with the Au thin
film resistor. Section IV B discusses the releasing of
patterned SU8 structures while Section IV C discusses the
procedures for establishing electrical interconnection to the
fabricated actuator device.
A. SU8 thermal actuator patterning
The process steps involved in the fabrication of SU8
microcantilever is illustrated in Fig. 4(a-e). (a) The process
begins with RCA cleaning of p-type {100} silicon wafer
followed by wet thermal oxidation in order to produce 1 µm
silicon dioxide layer on the silicon wafer. This oxide layer
serves as an adhesion layer to pattern SU8 and also as a
sacrificial layer for SU8 cantilever release. (b) 5nm/50nm
thick chromium and gold (Cr/Au) films are sputter deposited
and patterned to form contact pads over the oxide layer. (c)
Cr/Au resistors of 5nm/50nm thickness were sputtered and
patterned over the Cr/Au contact pads. During Cr/Au
Fig. 4. Schematic showing the sequence of fabrication.
B. Releasing patterned SU8 thermal actuator
The patterned SU8 microcantilevers were released by
immersing the silicon wafer in 25% buffered hydrofluoric
acid (BHF) for 4-5 hours. Since SU8 and Au do not react in
BHF, only the silicon layer oxide layer between the SU8
layer and the silicon wafer gets etched, thereby releasing the
SU8 structures from the silicon wafer. The released SU8
chips with suspended microcantilevers remain floating on the
surface of the aqueous BHF. These structures were carefully
removed using polymer tweezers. Fig. 5(a) & (b) shows a
fully released SU8 microcantilever with integrated Au thin
film resistor on its surface.
Fig. 5.(a)Scanning Electron Micrograph (SEM) of released SU8
microcantilever. (b) Optical microscope image of suspended SU8
microcantilever with metal thin film resistors.
An important issue with fabricating SU8 microcantilever
by surface micromachining technique is the influence of
residual stress on the shape of the suspended microstructure.
This stress occurs due to large differences in the CTE values
of sacrificial layer(silicon dioxide, 0.56ppm/ºC) and SU8
layer(52ppm/ºC). In addition to the poor adhesion of SU8
onto Au surface, the CTE mismatch between SU8 and silicon
dioxide induces a tensile residual stress on the SU8 cantilever
surface, thereby sometimes resulting in peeling-off and
cracking of Au thin film (Fig. 6(a)). Further, the evaporation
of solvents during the processing of SU8 negative photoresist
after UV exposure results in non-uniform stress across the
thickness of SU8 layer. This stress gradient induces an
upward curvature in the SU8 cantilever as seen in Fig. 6(b).
Also, Fig. 6(c) shows the cracks on SU8 films just after post
exposure baking and development. These film cracks and the
undesirable cantilever curvature can be minimized by
controlling SU8 soft baking and post exposure baking
temperatures during lithography [32].
bonding technique was adopted by combining Au ball bump
formation along with conductive epoxy bonding methods
[33]. First, a suitable printed circuit board (PCB) was
designed such that the Au contact pads on the fabricated SU8
chip are aligned with the PCB contact pads (Fig. 7(a)). Then,
Au ball bump was formed on the surface of PCB contact pad
by ball wedge bonding. The Au ball bump ensures electrical
connectivity between the SU8 chip and the PCB. Finally,
SU8 chip was attached to the PCB pads using silver
conductive epoxy (H20E EPO-TEK, Ted Pella, Inc.). The
cross-sectional view of Au ball bump on PCB with silver
epoxy underfill is seen in Fig. 7(b). Once the SU8 chip
attached, it is slightly pressed against PCB to planarize the
silver epoxy underfill and the epoxy is cured at room
temperature for 48 hours. The optical microscope image of
SU8 chip bonded over PCB is seen in Fig. 7(c). The silver
epoxy not only provides good electrical connectivity, but also
ensures better adhesion of SU8 chip with PCB.
The average electrical resistances of Au resistors, measured
just after flip-chip bonding was found to be about 120 Ω.
After connecting wires were soldered onto PCB with SU8
chip, the overall resistance was in the range 120-130 Ω.
Fig. 7. (a) Schematic showing a SU8 chip contact pads aligned with the
PCB’s. (b) Schematic cross-section showing Au ball bump with silver epoxy
underfill.(c) Optical image showing SU8 chip attached to PCB via flip-chip
Fig. 6. (a) SEM image showing peeling off Au thin film. (b) SEM image
showing cantilever curvature due to residual stress. (c) Optical image
showing the cracks on patterned SU8 layer
C. Electrical Interconnection
After fabrication of the SU8 based thermal actuators, it is
necessary to make electrical contact with them. Conventional
wire bonding techniques used in packaging silicon based
MEMS devices fail in the case of SU8 polymer material. This
is because polymers are good thermal insulators and also
absorb ultrasonic vibrations, thereby making conventional
wire bonding difficult. In this work, therefore, flip-chip
This section discusses the results of experimental
characterization of the actuator. Section V A describes the
set-up employed for performing characterization. Section V B
discusses the release of residual stress in the actuators.
Section V C discusses the experimentally measured
deformation profile of the actuator and compares them with
analytical results. Finally, Section V D discusses the possible
improvements that can be achieved in the analytical model by
incorporating nonlinear temperature dependence of the
thermal expansion of SU8 material.
A. Experimental setup for characterization of the SU8
thermal actuator
The experimental arrangement used to determine the
thermomechanical response of a fabricated SU8 thermal
actuator is illustrated in Fig. 8(a). Keithley source-meter
(2440 5A) was used to supply constant DC current to the Au
resistor in order to produce thermal actuation of SU8
cantilever. The subsequent SU8 cantilever deflection was
measured using non-contact 3D optical profiler (Talysurf,
CCI). A focused optical laser light was scanned over the
actuator and reference cantilever to obtain their relative
displacements along Z-axis. Fig. 8(b) shows the optical
microscope image of the fabricated actuator cantilever anlong
with the reference cantilever when the actuation current, I = 0
mA. Also, it’s 2D surface profile is seen in Fig. 8(c). Thermal
actuation experiment was carried out in air at room
temperature and the experimental observations are discussed
in the following section.
Fig. 8. (a) Schematic showing the experimental set-up used for the
characterization ofSU8 thermal actuator. (b) Optical microscope image of the
SU8 microcantilevers when I= 0 mA. (c) Typical 2D surface profile of the
actuator and reference SU8 microcantilevers obtained using 3D optical
profiler when I= 0 mA.
B. Release of residual stress in the SU8 thermal actuator
Residual stress in the actuator arises due to mismatch
between the CTEs of SU8 and the silicon dioxide substrate.
Some papers [32], [34] discuss the significance of fabrication
processing conditions such as hard bake temperature and
baking time in reducing the effect of residual stress during the
fabrication of SU8 microstructures. While there are several
charaterization techniques used to measure residual stress in
conventional silicon MEMS devices [35], there are few
reportson characterizing residual stress in polymeric MEMS
structures. The effect of residual stress can be observed
during the thermal actuation process. Fig. 9(a)shows the
thermal response of SU8 thermal actuator, measured with
respect to its evaluated by passing current from 0 to 5 mA in
steps of 0.5 mA. The plot labeled ‘Trial_1’ in Fig. 9(a) shows
the deflection of the actuator before release of thermal stress.
It is seen from the plot that the actuator deflects downwards,
i.e., in a direction opposite to the expected trend. However,
upon increasing the current beyond 4mA, for which the
maximum rise in temperature is 125ºC, the actuator starts to
deflect in the expected manner, presumably due to release of
thermal stress. A similar strategy, of gradually heating the
polymer close to its softening temperature, has also been
suggested [21]. Once the stress is released, the subsequent
actuation trials do not show the initial trend, but instead
deflect as expected. Further, the deflections were found to be
repeatable over several cycles of actuation (traces Trial_2,
Trial_3, and Trial_4 in Fig. 9(a)). Figs. 9 (b) and (c) show
the optical profilometer images of the profile of the actuator
before and after release of thermal stress respectively. The
reference cantilever seen in the 3D deflection profiles is used
to compare with the actuator deformation profile after
releasing residual stress.
Fig. 9. (a) Thermal response of SU8 thermal actuator illustrating residual
stress release during actuation process. (b) The position of actuator before
releasing the residual stress when actuation current, I=0. (c) Slightly bent
position of an actuator after releasing the residual stress when actuation
current, I= 4 mA.
It is worth noting that care has to be exercised in passing
appropriate actuation current during this process, so that the
maximum temperature does not exceed the softening
temperature of SU8 (225ºC). When the current exceeds 6
mA, the corresponding tempertaure rise in the actuator is
about 200ºC. At this temperature the SU8 polymer softens,
losses mechanical stability and undergoes permanent,
irreversible curling as seen in Fig. 10(b). However, this
curling effect can be avoided by operating the actuator well
below 6 mA.
Fig. 10.SEM image showing (a) SU8 microcantilever before thermal
heating.(b) The irreversible curling of SU8 microcantilever after 6mA
actuation current is passed. 40º tilted view of the curled up SU8 actuator is
C. Thermomechanical actuation of the SU8 thermal actuator
The thermomechanical performance of a stress-free SU8
actuator was experimentally obtained using 3D optical
profiler. Fig. 11(a-d) shows 3D profiles of SU8 actuator
deflection for different actuation currents, while Fig. 11 (e)
shows the corresponding 1D deflection profiles of the
actuator. The actuator displayed a small deflection even when
the current I through the actuator was zero, owing to the
release of thermal stress. The deflection of the actuator at
higher currents was measured relative to this initial profile.
When the current I was increased upto 5 mA, the rise in
temperature within the SU8-Au bimorph induces
corresponding thermomechanical deflection of the actuator as
seen in Fig. 11(b-d). When the actuation current is about 6
mA the deflection of the actuator exceeds the measurement
range, of the optical profiler. Fig. 11(f) compares the
experimentally obtained deflection profile of the actuator,
relative to its initial profile, at one particular current, viz., I=4
mA, with the results of non-linear thermomechanical
analysis. It is observed that the experimental deformation
profile agrees with the non-linear FEA with a maximum
difference between the two being about 14.3%. Similar
agreement between experiment and FEA were also found at
other actuation currents.
Fig. 11.(a-d) 3D profile showing SU8 thermal actuator deflection at different
actuation currents.(e) 1D cantilever deformation profile of the SU8 actuator.
(f) 1D deformation profile of the actuator obtained experimentally is
compared with non-linear FEA for an actuation current of 4mA.
The thermal response for a single heating and cooling cycle
of a stress-free actuator is shown in Fig. 12(a). The response
indicates low hysteresis and good repeatability. In order to
compare the experimentally observed thermal response of the
SU8 thermal actuator, Fig. 12(b) plots the experimentally
observed responses from three distinct actuators along with
the results of both nonlinear and linear FEA. It can be
observed that the three actuators show maximum deflection
about 6 to 10 μm for an actuation current of 5 mA. The
nonlinear FE simulation results are seen to agree well with
the experimentally observed thermal response, but slightly
underestimate the deflection, especially at higher currents. In
contrast, the results of linear FEA, which ignores the
variation of Young’s modulus and Poisson’s ratio with
temperature, is seen to underestimate the deflection by over
70%. This underscores the need to incorporate the
dependence of mechanical properties of SU8 on temperature
when analyzing the overall mechanical deflection of a SU8
based thermal actuator.
chip bonding technique was employed. The thermal response
of the actuators was experimentally evaluated. The fabricated
actuators showed the effects of residual stress and resulted in
anomalous deflection in the first trial. However, it was shown
that heating the actuator adequately can alleviate the effect of
residual stress. Subsequently, the actuator was demonstrated
to show repeatability and low hysteresis in response. The
experimentally measured average actuation range, of about
8.5μm for a current of about 5mA, was found to be in good
agreement with theory. It was demonstrated that
incorporation of the nonlinear thermal expansion of SU8 in
the model can further improve the correspondence between
experiment and theory. The analysis and experimental
evaluation results presented in this paper enable precise
design of thermally actuated micromanipulation tools for
robotic microgrippers and micro tweezers. They also provide
a strategy to eliminate residual stress after fabrication.
Fig.12. (a) Thermal response of a stress-free SU8 thermal actuator obtained
experimentally. (b) Comparison of experimental and FEA results.
The authors wish to acknowledge the Centre for Nano
Science and Engineering (CeNSE), Indian Institute of
Science, Bangalore for providing microfabrication and
Prof.M.M.Nayak of CeNSE for his kind suggestions and help
regarding the packaging aspects of the actuators.
D. Discussion
While the results of finite element analysis generally agree
with the experimental results, the difference between the two
can be attributed primarily to errors in estimation of the TEC
and the Young’s modulus of SU8. Further, it is noticed that
the error is significant primarily for larger currents, for which
case, the operating temperature of the actuator is close to the
softening temperature of SU8. Since it is reasonable to expect
small nonlinearity in CTE in this temperature range, it is
assumed that the CTE "2 " of SU8 demonstrates a variation
with temperature as,
 2 (T )   2 (T0 ) 1   T  T0  
Where, the constant 2 (0 ) is identified to be 52 ppm /ºC.
Then, the dashed curve in Fig. 12(b) shows the estimated
deflection profile when the value of ε = 0.00175ºC -1. It is
evident that the new curve is closer to the experimentally
observed deflection. Thus, incorporation of nonlinearity in
CTE, based on physical considerations of the operating
temperature, is seen to result in a significantly better
agreement between the nonlinear finite element model and
the experiment.
This paper reported design, thermomechanical analysis,
fabrication and evaluation of an SU8 based thermal actuator.
Thermomechanical analysis of the response of the actuator
was performed by incorporating the temperature-dependant
mechanical properties of SU8. Subsequently the actuator was
fabricated using surface micromachining techniques. In order
to make electrical connection with the polymer actuator, flip-
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