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ORIGINAL RESEARCH ARTICLE
published: 12 September 2013
doi: 10.3389/fneur.2013.00124
A highly compliant serpentine shaped polyimide
interconnect for front-end strain relief in chronic
neural implants
Viswanath Sankar1, Justin C. Sanchez2, Edward McCumiskey 3, Nagid Brown3, Curtis R.Taylor3,
Gregory J. Ehlert3, Henry A. Sodano3and Toshikazu Nishida1*
1Department of Electrical and Computer Engineering, University of Florida, Gainesville, FL, USA
2Department of Biomedical Engineering, University of Miami, Coral Gables, FL, USA
3Department of Mechanical and Aerospace Engineering, University of Florida, Gainesville, FL, USA
Edited by:
Nicholas Hatsopoulos,The University
of Chicago, USA
Reviewed by:
Jit Muthuswamy, Arizona State
University, USA
Alireza Mousavi, Brunel University, UK
*Correspondence:
Toshikazu Nishida, Department of
Electrical and Computer Engineering,
University of Florida, 219 Larsen Hall,
968 Center Drive, Gainesville, FL
32611-6200, USA
e-mail: nishida@ufl.edu
While the signal quality of recording neural electrodes is observed to degrade over time,
the degradation mechanisms are complex and less easily observable. Recording microelec-
trodes failures are attributed to different biological factors such as tissue encapsulation,
immune response, and disruption of blood-brain barrier (BBB) and non-biological factors
such as strain due to micromotion, insulation delamination, corrosion, and surface rough-
ness on the recording site (1–4). Strain due to brain micromotion is considered to be one
of the important abiotic factors contributing to the failure of the neural implants.To reduce
the forces exerted by the electrode on the brain, a high compliance 2D serpentine shaped
electrode cable was designed, simulated, and measured using polyimide as the substrate
material. Serpentine electrode cables were fabricated using MEMS microfabrication tech-
niques, and the prototypes were subjected to load tests to experimentally measure the
compliance. The compliance of the serpentine cable was numerically modeled and quan-
titatively measured to be up to 10 times higher than the compliance of a straight cable of
same dimensions and material.
Keywords: neuroprosthetics, brain-machine interface, flexible microelectrode array, strain relief, high compliance
electrode cable
INTRODUCTION
Recent clinical trials have successfully demonstrated that Brain-
Machine Interfaces (BMIs) can restore the lost communication
and control in humans affected with a variety of neurological
disorders (5–10). Though these studies have shown proof of con-
cept of BMI function, there is still the challenge of building an
ideal BMI system that is capable of obtaining high quality neural
signals for chronic durations (10+years), which is a desirable
requirement for clinical deployment. The temporal degradation
of signal quality in chronically implanted microelectrode neural
interfaces is attributed to both biotic factors such as tissue encap-
sulation, immune response,and disruption of blood-brain barrier
(BBB) and abiotic factors such as insulation delamination, corro-
sion, surface roughness on the recording site, and strain due to
micromotion (1–4).
Strain due to micromotion is identified as one of the poten-
tial abiotic contributors of failure mechanisms for long-term
neural implants. Histological studies (11,12) report that the strain
induced immune response caused by the rigid tethering of the
electrode to the skull showed an increase in microglial activity
in the implanted tissue as compared to untethered electrodes.
This increased tissue response and continuous proliferation of
microglial cells around the electrode can be detrimental to the well
being of the neurons in the vicinity. Histological studies (13) report
that upregulation of microglial biomarker ED1 was accompanied
by reduction in neurons and nerve cell fibers surrounding the
implant. This suggests a correlation between increased tissue
response and reduced signal reliability.
When the electrode substrate is secured to the skull during
implantation, it results in a rigid tethering of one end of the elec-
trode, while the other end of the electrode, the tip, and the brain
are free to move with respect to each other. The brain experi-
ences displacements on the order of microns to millimeters driven
by physiological, behavioral, and mechanical sources (14). It has
been observed that the brain micromotion in anesthetized rats
due to respiratory pulsation is on the order of 10–30 µm, and due
to vascular pulsation is about 2–4 µm (15). This is relevant since
micromotion of the brain with respect to the skull (relative micro-
motion) is expected to exert a force vector on the cortical tissue
via the implanted electrode. The strain due to the force acting on
the rigid back end is transferred along the probe shank and dis-
places the electrode tip within the brain tissue and may have biotic
ramifications through a mechanical inflammatory process with
consequences such as promoting more upregulation of microglial
cells.
Numerical studies have shown that electrodes with lowYoung’s
modulus material or redefined geometry for high compliance can
provide front-end strain relief. Mechanical modeling of tethering
induced strain for silicon and polymer electrodes (16) show that
a rigidly tethered silicon shank transfers significant strain to the
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Sankar et al. Highly compliant serpentine shaped interconnect
surrounding brain tissue and favor displacement of tip. Finite ele-
ment analysis of a polyimide array with respect to a rigid silicon
microelectrode array of same dimensions conducted (17,18) show
a front-end strain relief of 65–94%.
Polymers such as polyimide and parylene-C, with their low
Young’s modulus values and good biocompatibility have been the
choice of researchers for electrode substrate material. In the past
decade, a number of groups have developed microelectrode arrays
with polyimide as the electrode structural material (19–22), and
parylene-C as the structural material (23,24).
In terms of design modification for front-end strain relief, the
first effort (25) included a single long slender gold wire as a cable
between the probe and the external connection. The design limited
the electrode displacement to 10 µm for 1 mm relative micromo-
tion of the brain. Of the several electrode array configurations
reported by (19), one included an“S”-shaped curve for incorporat-
ing strain relief in the cable. Recently,the serpentine shaped silicon
cable showed 50% stress reduction compared to a straight silicon
cable of the same dimension (26). More recently, neural probes and
cables with different meandering geometries are being reported
with improved compliance and better stress relief (27,28).
In this paper, we describe the design, fabrication, and detailed
mechanical characterization and numerical studies of highly com-
pliant serpentine shaped polyimide modular electrodes. Modular
electrodes are the fourth generation polyimide electrodes devel-
oped at the University of Florida after the generation 1 flexible
polymer substrate electrodes (29), the generation 2 amplifier inte-
grated microelectrode arrays (30), and the generation 3 Pyrex®
supported amplifier integrated microelectrode arrays. These elec-
trodes have two modules: (i) a rigid silicon platform serving
as a stage for the connector and future electronics, and (ii) a
serpentine shaped flexible polyimide cable that interfaces the
recording tungsten microwire array and the rigid module. The
serpentine shaped cable possesses a higher compliance than a
straight polyimide cable, which can provide better front-end strain
relief.
MATERIALS AND METHODS
FLEXIBLE CABLE AND RIGID MODULE DESIGN
The modular electrodes include two modules – a rigid silicon
module serving as a platform for electronics and connector, and a
flexible polyimide module serving as the front-end cable between
the microwires and the electronics. The modules may be fabricated
independently and then bonded together using conductive silver
epoxy paste, and secured and hermetically sealed with underfill
epoxy. This kind of modular approach allows parallel fabrica-
tion of the modules, thereby reducing the processing time and
increasing the yield. Figure 1 shows the conceptual drawing of the
modular electrode design.
The flexible polyimide cable module is designed to have a ser-
pentine structure as shown in Figure 1. The meanders in the design
result in an increase in the effective length comparedw ith a straight
electrode of the same overall size. This provides higher compliance
and better strain relief than its straight counterpart. At the same
time, the overall form factor of the cable is still maintained the
same for facilitating implantation. Furthermore, the new geome-
try enables placing the recording microwires in a two dimensional
transverse fashion, with nine electrodes being placed in a 3 ×3
array.
FLEXIBLE CABLE AND RIGID MODULE FABRICATION
Processing and packaging steps involved in the fabrication of the
2D cables are shown in Figure 2. All the steps involved in the pro-
cessing are done on a rigid 400 silicon wafer. The process begins
with the spin deposition of 20 µm thick layer of polyimide on top
of a sacrificial aluminum film. Next, a thin film (∼2000 Å) of gold
is sputter deposited and lithographically patterned to define the
conductive traces. The top insulation is provided by spinning a
layer of 20 µm thick polyimide over patterned gold. The top poly-
imide is plasma etched using a reactive ion etcher (RIE) to expose
gold bondpads surrounding the via holes. Also the through holes
are obtained by etching off bottom polyimide underneath them.
FIGURE 1 | Illustration of the modular electrode design. Image not
drawn to scale.
FIGURE 2 | Process flow for 2D transverse flexible cable module.
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Sankar et al. Highly compliant serpentine shaped interconnect
Finally the device is released from the Si wafer by etching the
sacrificial aluminum layer through anodic dissolution process.
The processing and packaging steps involved in the fabrica-
tion of the rigid electronics and connector module are shown in
Figure 3. All the steps are carried out on a 400 silicon wafer. First,
a thin film of SiO2is deposited on silicon. Similar to the process
flow of the flexible module, a thin film (∼2000 Å) of gold is sput-
ter deposited on the oxide layer and lithographically patterned
to define the conductive traces. The top insulation is deposited
by spinning a layer of 20 µm thick polyimide over the patterned
gold. The top polyimide is plasma etched using RIE to expose gold
bondpads for the connector and the electronics. The hole for the
ground screw and the slot hole are obtained by etching off the
polyimide and oxide underneath them. Finally, the silicon under-
neath the screw hole, slot hole, and the surrounding device outline
is etched off using deep reactive ion etch (DRIE) to separate the
module. Once the modules are separated into single pieces, the
connector and the nut are attached using conductive silver epoxy.
Photographs of the fabricated serpentine cable prior to
attachment of the microwires and rigid module are shown in
Figures 4A,B, respectively.
ANALYTICAL AND NUMERICAL MODELING OF CABLE COMPLIANCE
Analytical analysis
After implantation, one end of the electrode cable is attached
firmly to the skull resulting in a rigid tethering, while the other end
containing the recording microelectrode array is inserted into the
cortex, as shown in Figure 5A. The natural respiration and vascular
FIGURE 3 | Process flow for rigid electronics and connector module.
pulsation along with head movement of the animal may result in
the micromotion of the brain with respect to the skull. The brain
exhibits motion in all the three axes. The translational movement
of brain with respect to the skull in xand yaxes exerts a force both
radial and tangential to the electrode cable (17). In addition, the
translational movement of the brain in the axis orthogonal to the
plane of the electrode (z-axis) contributes to the shearing of the
cable. However, for this analysis,three assumptions were made:
1. the rotational motion of the brain and hence the torsion of the
electrode is not considered,
2. the force acting on the electrode is assumed to be a point load
concentrated at the cortex end of the electrode cable, and
3. the cable is fixed in the vertical direction at the surface of the
skull.
Guided by these boundary conditions and assumptions, a beam
model was employed for the electrode cable. According to the beam
model, the cable is considered as a clamped-guided beam, which
is fixed at the skull and free to move in x,y, and zaxes along
the cortex, and a concentrated force is acting at the guided end.
Figures 5B,C show the free-body diagrams of the straight cable
and serpentine cable respectively.
Closed-form theoretical expressions for classic serpentine
springs (31) were used for calculating the spring constants and
compliances of the serpentine cable in x,y, and zdirections.
FIGURE 4 | (A) Photograph of the micro-fabricated serpentine polyimide
cable prior to the attachment of the tungsten microwires and
(B) photograph showing the fabricated silicon rigid module packaged with
the Omnetics connector and the ground screw compared against a one
cent coin.
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Sankar et al. Highly compliant serpentine shaped interconnect
FIGURE 5 | (A) Schematic drawing of the electrode implanted into the brain, (B) free-body diagram of the straight cable, and (C) free-body diagram of the
serpentine cable. Images not drawn to scale.
The stiffness equations in the x,y, and zaxes for the serpentine
cable are given by (31),
KC
x="(N+1)l3
o
6EIzo+(N+1)l2
olp
2EIzp#−1
, (1)
where Nis the number of meanders, lois the length of the mean-
der element perpendicular to the xand zaxes (m), lpis the length
of the meander element parallel to the x-axis, Izois the moment of
inertia with reference to the z-axis of the meander element section
perpendicular to xand zaxes (m4), and Izpis the moment of iner-
tia with reference to the z-axis of the meander element section
parallel to x-axis (m4).
KC
δy=
KC
yθzKC
θzyKC
y
KyθC
zKC
θzy−KC
yKC
θz
, (2)
where,
KC
y="2(N+2)lp3
3EIzp+8N3+36N2+55N+27l2
plo
3EIzo#−1
,
(3)
KC
yθz=KC
θzy ="2N2+3N+4lplo
EIzo+(2N+2)2l2
p
EIzp#−1
,
(4)
KC
θz="2(N+2)lp
EIzp+2(N+1)lo
EIzo#−1
, (5)
and,
KC
δz=
KC
zθyKC
θyzKC
z
KC
zθyKθyzC−KC
zKC
θy
, (6)
where,
KC
z="2(N+2)lp3
3EIyp+(N+1) (lo)3
6EIyo+(N+1) (lo)2lp
GJp
+8N3+36N2+55N+27l2
plo
3GJo#−1
, (7)
KC
θyz=KC
zθy"2N2+3N+4lplo
GJo+2(N+2)2l2
p
EIyp#−1
, (8)
KC
θy="2(N+2)lp
EIyp+2(N+1)lo
GJo#−1
, (9)
where Iyois the moment of inertia with reference to the y-axis
of the meander element section perpendicular to xand zaxes
(m4), and Iypis the moment of inertia with reference to the y-axis
of the meander element section parallel to x-axis (m4), Jois the
cross-sectional torsion factor of the meander element perpendic-
ular to xand zaxes (m4), Jpis the cross-sectional torsion factor
of the meander element parallel to x-axis (m4), and Gis the shear
modulus (Pa).
The spring constants for the straight cables were calculated
using the standard stiffness equations for a clamped-guided beam
(32). The stiffness equations in x,y, and zaxes are given by,
Kx=Ehw
L, (10)
Ky=Ehw3
L3, (11)
and
Kz=Eh3w
L3(12)
Since the stress due to complex mechanical motions and biolog-
ical and chemical reactions are not considered for this analysis,the
standard stiffness equations may be used to estimate the straight
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Sankar et al. Highly compliant serpentine shaped interconnect
cable compliance. However, it should be noted that in a real time
in vivo condition, these equations may not completely model the
cable compliance.
Table 1 gives the dimensions of the straight and the serpen-
tine cables used in the analytical and numerical estimation of
compliance.
Material uncertainty. In the literature, the reported Young’s
modulus of polyimide ranges from 2.793 to 15 GPa depending
on the formulation. Table 2 gives a list of all Young’s modulus
values for polyimide obtained from the literature.
The compliance was calculated for all of the values listed in
Table 2 for both the straight and the serpentine cables. Mean
and standard deviations were obtained for the calculated com-
pliance values to account for the error due to variation in material
property.
Numerical analysis
Following the analytical analysis, a numerical analysis was per-
formed. A finite element model was developed for the compliances
of straight and serpentine beams using ABAQUS simulation tool.
The 2D Shell element was used for the analysis. The mesh shape
is quadrilateral element for straight cables and quadrilateral and
triangular elements for serpentine cables. The mesh size is within
the ABAQUS default size range (min: 1.3e−6 and max: 0.0013),
and the mesh number is 50 elements for straight cables and 833
elements for serpentine cables. The test for convergence employs
the ABAQUS default convergence criteria values for non-linear
problems as described in ABAQUS Analysis User’s Manual1. The
boundary conditions used were encastré (no rotational or transla-
tional motion in any axis) for fixed end and no rotational motion,
only translational displacement in all axes for the guided end. Sim-
ulations were performed for all of the Young’s modulus values of
polyimide given in Table 2 and the results were averaged and the
error was calculated.
EXPERIMENTAL MEASUREMENT CABLE COMPLIANCE
Experimental measurement of the in-plane (x-axis) cable
compliance
The in-plane (or x-axis) stiffness of the micro-fabricated straight
and serpentine cable electrodes was experimentally measured
using an Instron® 5900 series mechanical testing system. The
cables were subjected to in-plane tensile stress, and the extension
in the axial direction was measured to evaluate the stiffness in the
x-axis. The naming of the axes can be found in Figure 5.
The cables were suspended between two vertically positioned
clips that were connected to load cells, which run across two load
frames. The top frame is movable while the bottom remains fixed.
Extension control testing method was used to measure the compli-
ance of the cables. In this testing method, the cable is subjected to a
known extension (from x0to x1in steps of ∆x) for a given period
of time. From Hooke’s law, F=kx, a change in displacement is
induced by a change in the force. The corresponding change in
the force ∆Fis measured. The measured change in force is plotted
1http://abaqus.me.chalmers.se/v6.12/books/usb/default.htm?startat=
pt03ch07s02aus51.html
Table 1 | Dimensions of the straight and serpentine cables used in the
analytical and numerical analysis.
Dimension Straight cable Serpentine cable
Length 6 mm 6 mm
Width 2.7 mm 2.5 mm
Height 40 µm 40 µm
Number of meanders – 2
Length of meander – 1 mm
Width of meander – 1 mm
Length of end spans – 0.5 mm
Table 2 | List ofYoung’s modulus values of polyimide obtained from
literature.
Reference Young’s modulus
of polyimide (GPa)
Rousche et al. (19) 2.793
Dupont Kapton B technical datasheeta3.0
Dolbow and Gosz (33) 7.5
HD microsystems PI2611 process guideb8.5
Dolbow and Gosz (33) 8–15
ahttp:// www2.dupont.com/ Kapton/ enUS/ assets/ downloads/ pdf/ KaptonB.pdf
bhttp://hdmicrosystems.com/HDMicroSystems/enUS/pdf/PI-2600Process
Guide.pdf
against the change in displacement, and the slope of the resulting
curve is calculated to find the stiffness of the cable.
In order to ensure the consistency of cable length with the ana-
lytical and numerical calculations, the cables were mounted on
paper tabs with holes in the center using hot glue or crystal bond.
The size of the holes corresponded to the length of the cable.
The paper tabs were then attached to the two vertical clips. Once
attached, the paper was cut at the center to prevent any additional
loading due to the paper. This set up has more control on the gage
length of the cable. Figure 6A shows the schematic drawing of the
paper mounted cable suspended between the two vertical clips,
and Figure 6B shows a photograph of the cable mounted on the
paper tab connected to the clips of the Instron® mechanical testing
system.
The straight cables were extended from 0 to 50µm. The max-
imum load limit set on the straight cables was 300 mN. The
serpentine cables were extended from 0 to 40µm, and the max-
imum load limit on them was 100 mN. One sample of straight
cable and one sample of serpentine cable were measured for stiff-
ness and 10 trials were performed on each sample for consistency.
Mechanical deformation such as necking was not observed in
the in-plane stiffness measurement experiments as the load was
applied within the elastic region. Also no buckling was observed
since no compressive stress was applied.
Experimental measurement of the out-of-plane (y-axis) cable
compliance
The out-of-plane (or y-axis) stiffness of the micro-fabricated
straight and serpentine cable electrodes was experimentally
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Sankar et al. Highly compliant serpentine shaped interconnect
FIGURE 6 | (A) Schematic diagram of the paper mounted cable suspended between the two vertical clips connected to the load cells and (B) photograph of the
cable mounted on the paper tab connected to the clips of the Instron® mechanical testing system. Inset shows the closer view of the mounted cable.
measured using a Hysitron® TriboIndenter (TI 900). The instru-
ment has a lab noise floor displacement resolution of 1 nm, and
force resolution of 100 nN. The cables were loaded by lowering
the indenter tip down for a preset load limit and a depth limit,
and the change in the displacement was measured as a function of
the change in the load to evaluate the stiffness. A diamond (fluid
cell) conical tip of nominal radius 20.1 µm was used for load-
ing the samples. The loading conditions included a preload force
of 0.3 µN with a triangular ramp of 100nm/s and a maximum
displacement of 4.8 µm.
Sample preparation. In order to facilitate vertical loading of the
cables along its thin y-axis, the cables were mounted vertically on
a glass slide and secured with two magnet pieces and steel nuts
on either side. This set up ensured vertical standing of the cable
without much movement. It was also confirmed that the nuts and
magnets remain steady during the loading of the tip. One sample
of straight cable, and one sample of serpentine cable were prepared
for the experiment.
EXPERIMENT SETUP
The mounted sample was introduced into the TriboIndenter
chamber, and placed directly under the conical tip. The system
was allowed to thermally stabilize for 10min. First, air indent cal-
ibration was done. Next, the initial position (or the zero point)
was calibrated by slowly lowering the tip and establishing a con-
tact with the cable at a force <2µN. After calibration, the cable
was loaded by lowering the tip further down for a preset load and
depth limit. The maximum load limit was set at 25 µN, and the
maximum depth was set at 4700 nm. The change in the displace-
ment was measured as a function of the change in the load, and the
slope of the resulting plot was calculated to obtain the compliance
value. Five trials were performed on each sample for consistency
and statistics. Figure 7 shows the schematic drawing of the vertical
loading of the serpentine cable in y-axis by the conical nanoinden-
ter tip along with a photograph. The y-axis stiffness measurement
FIGURE 7 | Schematic diagram showing the vertical loading of the
serpentine cable in y-axis by the nanoindenter tip. Inset showing the
photograph of the cable and the nanoindenter tip.
experiments were conducted within the elastic region of the cables
and hence no tearing due to fracture was observed. Further, since
the gold traces are very thin (200 nm) when compared to the
polyimide film (∼40 µm thick), their contribution to the over-
all bending stiffness will be <1%, and hence can be neglected. The
indent test was conducted at room temperature and an in situ
optical microscope was used to align the tip with the location of
indent on the sample.
Experimental measurement of the out-of-plane (z-axis) cable
compliance
The out-of-plane (or z-axis) stiffness of the micro-fabricated
straight and serpentine cable electrodes was experimentally mea-
sured using a Hysitron® TriboIndenter (TI 900). The cables were
loaded by lowering the indenter tip down for a preset load limit and
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Sankar et al. Highly compliant serpentine shaped interconnect
a depth limit, and the change in the displacement was measured
as a function of the change in the load to evaluate the stiffness.
To minimize experimental errors, the same conical tip of nominal
radius 20.1 µm used in y-axis measurement was used for loading
the samples. Similar to y-axis experiments, the loading condi-
tions included a preload force of 0.3 µN with a triangular ramp of
100 nm/s and a maximum displacement of 4.8 µm.
Sample preparation. Prior to the experiment, the electrode sam-
ples were prepared to facilitate the loading of the nanoindenter tip
on the edge of the cable end without any damage to the electrode
and the tip. The electrode substrates were mounted on a glass slide
and firmly secured with hot glue at the back end. The cable end
of the substrate was extended beyond the edge of the glass slide to
allow free movement upon loading. A thin microscope slide was
added on top of the back end to ensure flatness of the substrate.
The mounted samples were then placed on the bottom plate of the
nanoindenter. Since the cables are loaded vertically by the nanoin-
denter tip, it is required to have enough clearance in the plane of
loading. Two small pieces of magnets were placed underneath the
glass slide to increase the clearance of the sample in the z-axis.
One sample of straight cable and one sample of ser pentine cable
were prepared for the experiment. After completing the experi-
ment, the samples were removed from the glass slides by heating
the glue and separating the slides.
Experimental setup. Sample mounted on the bottom plate was
introduced into the chamber of the Hysitron® TriboIndenter, and
placed directly under the conical tip. At first, air indent calibration
was done similar to y-axis loading. Later, the zero point was cali-
brated by slowly lowering the tip and establishing a contact with
the cable at a force <2µN. After calibration, the cable was loaded
by lowering the tip further down for a preset load limit and a depth
limit. The maximum load limit was set at 25 µN and the maxi-
mum depth was set at 4700 nm. The change in the displacement
was measured as a function of the change in the load, and the slope
of the resulting plot was calculated to obtain the compliance value.
The experiment was conducted at three different points on the
cable – the left most tip, the right most tip, and the center point.
Five trials were performed on each measuring point for consis-
tency and statistics. The data obtained for the right and left tips
showed high non-linearity for serpentine cable. Hence only the
data obtained at the center point, which were linear, were used
for analysis. Figure 8 shows the schematic drawing of the vertical
loading of the cable in the z-axis by the nanoindenter tip and a
photograph of the nanoindenter tip and the cable. The z-axis stiff-
ness measurement experiments were conducted within the elastic
region of the cables and hence no tearing due to fracture was
observed. Further, since the gold traces are very thin (200 nm)
when compared to the polyimide film (∼40 µm thick), their con-
tribution to the overall bending stiffness will be <1%, and hence
can be neglected. The indent test was conducted at room temper-
ature and an in situ optical microscope was used to align tip with
location of indent on the sample.
Uncertainty analysis of the experimental error
The instrument error was included for in-plane experiments
by calculating the variability in the load measurement for each
FIGURE 8 | Schematic diagram showing the vertical loading of the
cable in the z-axis by the nanoindenter tip. Inset shows the zoomed in
photograph of the tip and the cable.
measured reading based on the accuracy values (±5%) obtained
from the instrument manual2. The instrument error was included
for out-of-plane experiments by calculating the drift in trans-
ducer displacement for the loading duration. The value used for
transducer displacement drift was obtained from the instrument
manual as 0.05 nm/s3.
The measured values from different trials for each sample were
analyzed for experimental uncertainty and the confidence interval
was constructed. Two tailed t-test was used for confidence interval
calculations since the number of trials (samples) is <30.
The confidence interval for a t-distribution is given by,
C.I. = ¯x±tα/2, n−1s
√n, (13)
where, ¯x=mean, α=significance level (0.05 for 95% C.I.),
n−1=number of degrees of freedom, n=sample size, and
s=sample standard deviation, which is given by,
s=rs
n−1X
i=1n(xi− ¯x)2, (14)
RESULTS
Table 3 gives a summary of straight cable compliance estimated
using different methods for all three axes,and Table 4 gives a sum-
mary of serpentine cable compliance estimated using different
methods for all three axes. It can be observed (from the experi-
mental results) that the compliance value of the serpentine shaped
cables is at least one order of magnitude higher than the compli-
ance of the straight cables of the same dimensions. The higher
compliance or flexibility of the new serpentine shaped cables is
expected to lessen the front-end strain of the electrode on the
tissue. Mitigated front-end strain is expected to reduce the tis-
sue immune response and improve the reliability of the implant’s
signal recording quality.
25960 Series: Dual Column Tabletop (5–50 kN) manual,Instron Inc
3TI 900 TriboIndenter manual, Hysitron Inc
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Sankar et al. Highly compliant serpentine shaped interconnect
The measured compliance values for the straight and the ser-
pentine cables for x,y, and zaxes are given in Figure 9. It can be
noted from Figure 9A that the in-plane compliance values for the
serpentine cable is nearly 10 times higher than that for the straight
cable. Similarly, it can be observed from Figure 9B that the y-axis
compliance of the serpentine cable is nearly one order of magni-
tude higher than that of straight cable. Also from Figure 9C, it is
apparent that the serpentine cable has a compliance that is at least
six times higher than that of the straight cable in the z-axis.
In addition to these experiments, the compliance of the straight
and serpentine cables of reduced lengths (<6 mm) and increased
widths (≥2.5 mm) were calculated using analytical methods to
study the effect of variability in dimension on cable compli-
ance. The lengths used for the analysis were 4.2, 3.84, 3.8, and
3.44 mm, and the widths used were 2.5, 3,and 3.5 mm. T heYoung’s
modulus value of polyimide used for the analysis was 2.8 GPa.
All assumptions were same as described earlier. The analytical
Table 3 | Matrix comparing the straight cable compliance estimated
through analytical, numerical, and experimental analysis.
Type of analysis x-Axis
compliance
(m/N)
y-Axis
compliance
(m/N)
z-Axis
compliance
(m/N)
Analytical 1.12±0.67 ×10−55.54 ±3.29 ×10−50.252 ±0.15
Numerical 1.12±0.67 ×10−52.57 ±1.52 ×10−40.25±0.149
Experimental 5.69 ±0.48 ×10−57.43 ±1.07 ×10−30.188 ±0.06
Table 4 | Matrix comparing the serpentine cable compliance estimated
through analytical, numerical, and experimental analysis.
Type of analysis x-Axis
compliance
(m/N)
y-Axis
compliance
(m/N)
z-Axis
compliance
(m/N)
Analytical 1.89 ±1.12 ×10−32.7±1.6 ×10−226.6 ±15.78
Numerical 1.92 ±1.14 ×10−39.8 ±5.8 ×10−31.17 ±0.69
Experimental 5.34 ±0.19 ×10−46.8 ±1.8 ×10−21.54 ±0.56
results showed that the compliance of the serpentine cables was
much higher than the straight cables in all three axes, with high-
est increase in the y-axis. Further, the compliance increased with
the decrease in the cable length and increase in the cable width.
The analytically calculated compliance values for straight and ser-
pentine cables of reduced length and increased width are given in
Table 5.
DISCUSSION
We have developed a serpentine shaped polyimide electrode cable
that is nearly one order of magnitude more compliant than a
straight polyimide cable of same dimensions. The design para-
meters were carefully evaluated using analytical and numerical
models. Prototypes of the cables were micro-fabricated using
MEMS microfabrication techniques. Fabricated prototypes were
subjected to in-plane and out-of-plane stress experiments and the
compliance was measured. Measured compliance of the serpentine
cable was 6–10 times higher than the compliance of the straight
cable.
It is expected that a 10-fold increase in electrode interconnect
compliance may enable more reliable chronic neural recording
even from awake animals and the electrophysiological recordings
are expected to have better signal-to-noise ratio and impedance.
However, in vivo experiments are needed to validate and quantify
this. Though increased compliance in all three axes will be useful
for reduced stain, it is important to have more compliance in the
y-axis since the relative displacement of the brain with respect to
the skull will be highest in that direction (16). The new serpen-
tine cables are shown to be 10 times more compliant in the y-axis,
which will be critical in in vivo studies.
It should be noted that the Young’s modulus of polyimide
(3–15 GPa) is much less than that of silicon (200 GPa). With
one order of magnitude less Young’s modulus, the compliance
of the polyimide cables is expected to be much higher than
silicon ribbon cables. Furthermore, the serpentine structure pro-
vides additional flexibility to the interconnect. However,polyimide
has a water absorption rate of ∼1.1%. In vitro study of poly-
imide as a long-term implantable material has shown that poly-
imide of different formulations undergo changes in crystalline
A B C
FIGURE 9 | (A) Measured x-axis compliance values from straight and
serpentine cables. Load measurement variability was nearly negligible
(two orders of magnitude less) for straight and serpentine cables,
(B) measured y-axis compliance values from straight and serpentine
cables. The error bars represent the variability due to transducer
displacement drift during the loading period, and (C) measured z-axis
compliance values from straight and serpentine cables. Transducer
displacement drift during the loading period was nearly negligible (two
orders of magnitude less) for straight and (four orders of magnitude
less) for serpentine cables.
Frontiers in Neurology | Neuroprosthetics September 2013 | Volume 4 | Article 124 | 8
Sankar et al. Highly compliant serpentine shaped interconnect
Table 5 | Matrix comparing the analytically calculated compliances for short and wide straight and serpentine cables (overall cable length <6mm and overall cable width ≥2.5 mm).
Overall cable
width (mm)
Overall cable
length (mm)
x-Axis y-Axis z-Axis
Straight
cable
compliance
(m/N)
Serpentine
cable
compliance
(m/N)
Increase in
compliance of
serpentine
cable (times)
Straight
cable
compliance
(m/N)
Serpentine
cable
compliance
(m/N)
Increase in
compliance of
serpentine
cable (times)
Straight
cable
compliance
(m/N)
Serpentine
cable
compliance
(m/N)
Increase in
compliance of
serpentine
cable (times)
2.5 4.2 1.50 ×10−56.29 ×10−3419 4.24 ×10−50.034 801 0.166 24.57 148
3.84 1.37 ×10−58.59 ×10−3627 3.24 ×10−50.032 987 0.127 20.79 163
3.8 1.36 ×10−56.04 ×10−3444 3.14 ×10−50.027 860 0.122 20.92 171
3.44 1.23 ×10−58.24 ×10−3670 2.33 ×10−50.025 1072 0.091 17.45 191
3 4.2 1.25×10−59.62 ×10−3769 2.46 ×10−50.028 1138 0.138 25.58 185
3.84 1.15×10−51.32 ×10−21148 1.88 ×10−50.025 1329 0.105 21.83 208
3.8 1.13×10−59.28 ×10−3821 1.82 ×10−50.021 1154 0.102 21.88 214
3.44 1.03 ×10−51.27 ×10−21233 1.35 ×10−50.018 1333 0.076 18.42 242
3.5 4.2 1.07 ×10−50.014 1308 1.55 ×10−50.021 1355 0.118 26.73 226
3.84 9.82 ×10−60.019 1935 1.18 ×10−50.018 1525 0.09 23.04 256
3.8 9.72 ×10−60.013 1337 1.15 ×10−50.016 1391 0.087 23.04 265
3.44 8.79 ×10−60.018 2048 8.5 ×10−60.012 1412 0.065 19.56 301
www.frontiersin.org September 2013 | Volume 4 | Article 124 | 9
Sankar et al. Highly compliant serpentine shaped interconnect
structure and mechanical properties due to constant water uptake
(34). The changes in mechanical properties include increase in
Young’s modulus and decrease in tensile strength. The results
from the study predict that the constant uptake of water for
over 20 months could act as a plasticizer and can increase the
stiffness of the material. Based on these observations, it can
be implied that the compliance of the cables will decrease in
chronic conditions due to constant moisture uptake. Similarly,
the cables will undergo additional stress due to biological and
chemical reactions, and other complex mechanical motions in
in vivo conditions, which may potentially affect the compliance. It
is expected that the serpentine geometry will compensate for any
increase in the cable stiffness. Further studies are needed to eval-
uate the chronic in vivo behavior of these polyimide serpentine
cables.
It can be noted from Tables 3 and 4that there is some dis-
crepancy between the analytical and the numerical results and
the experimentally measured values. This discrepancy could have
resulted from the assumptions made for the analytical and numer-
ical analyses and from the limitations in the geometry of the finite
element solver. A clamped-guided beam model was assumed for
the analytical and the numerical analyses. The boundary condi-
tions of this model allow the guided end to deflect normal to its
axis, while restricting its rotational motion. However in practice,
there will be some rotational motion displayed by the beam which
will contribute to the overall beam deflection. Furthermore, in the
case of the serpentine structure, the meanders will have flexural
degrees of freedom which will be different from the rigid body
degrees of freedom, and there will be an additional effect of beam
twisting seen in the serpentine cables.
In addition to its compliance, the proposed cable design also
permits the placement of the recording microwires in a transverse
fashion thereby providing a 2D recording space for the array. The
2D high compliance serpentine electrode arrays are expected to
provide strain relief for the recording microwires and in turn, mit-
igate the strain induced tissue response. With a reduction in the
tissue response, the implant can be expected to have improved
performance in chronic applications.
ACKNOWLEDGMENTS
We thank the staff at the Nanoscience Research Facility (NRF) at
the University of Florida (UF) for their help with the fabrication of
the electrode cables. Also we thank the staff at the Major Analytical
Instrumentation Center (MAIC) at UF for their help with the com-
pliance experiments. Funding: the work was funded by NIH under
Grant number NS053561 and by the Defense Advanced Research
Projects Agency (DARPA) Microsystems Technology Office under
the auspices of Dr. Jack Judy through the Space and NavalWarfare
Systems Center, Pacific grant no N66001-11-1-4009.
SUPPLEMENTARY MATERIAL
The Supplementary Material for this article can be found
online at http://www.frontiersin.org/Neuroprosthetics/10.3389/
fneur.2013.00124/abstract
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Conflict of Interest Statement: The
authors declare that the research was
conducted in the absence of any com-
mercial or financial relationships that
could be construed as a potential con-
flict of interest.
Received: 09 May 2013; accepted: 19
August 2013; published online: 12 Sep-
tember 2013.
Citation: Sankar V, Sanchez JC,
McCumiskey E, Brown N, Taylor CR,
Ehlert GJ, Sodano HA and Nishida
T (2013) A highly compliant serpen-
tine shaped polyimide interconnect
for front-end strain relief in chronic
neural implants. Front. Neurol. 4:124.
doi: 10.3389/fneur.2013.00124
This article was submitted to Neuropros-
thetics, a section of the journal Frontiers
in Neurology.
Copyright © 2013 Sankar , Sanchez,
McCumiskey, Brown, Taylor, Ehlert ,
Sodano and Nishida. This is an open-
access article distributed under the terms
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provided the original author(s) or licensor
are credited and that the original publica-
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