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HomeScience AdvancesVol. 4, No. 3Stretchable ultrasonic transducer arrays for
three-dimensional imaging on complex surfaces
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STRETCHABLE ULTRASONIC TRANSDUCER ARRAYS FOR THREE-DIMENSIONAL IMAGING ON
COMPLEX SURFACES
Hongjie Hu https://orcid.org/0000-0002-3792-4894, Xuan Zhu, [...] , Chonghe Wang
https://orcid.org/0000-0002-1083-2666, Lin Zhang
https://orcid.org/0000-0002-0230-5911, [...] , Xiaoshi Li, Seunghyun Lee
https://orcid.org/0000-0002-1354-3261, Zhenlong Huang
https://orcid.org/0000-0002-3410-1411, Ruimin Chen
https://orcid.org/0000-0002-5338-359X, Zeyu Chen, [...] , Chunfeng Wang
https://orcid.org/0000-0001-9676-8425, Yue Gu
https://orcid.org/0000-0003-0437-5339, Yimu Chen, Yusheng Lei, Tianjiao Zhang
https://orcid.org/0000-0002-9758-4895, NamHeon Kim
https://orcid.org/0000-0002-3614-2233, Yuxuan Guo
https://orcid.org/0000-0001-7582-4555, Yue Teng
https://orcid.org/0000-0002-3032-3834, Wenbo Zhou
https://orcid.org/0000-0001-5980-3287, Yang Li, Akihiro Nomoto, Simone Sternini
https://orcid.org/0000-0002-6511-2666, Qifa Zhou, Matt Pharr
https://orcid.org/0000-0001-8738-5393, Francesco Lanza di Scalea, and Sheng Xu
https://orcid.org/0000-0002-3120-4992 shengxu@ucsd.edu+22 authors +20 authors
+15 authors fewerAuthors Info & Affiliations
Science Advances
23 Mar 2018
Vol 4, Issue 3
DOI: 10.1126/sciadv.aar3979
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ABSTRACT
Ultrasonic imaging has been implemented as a powerful tool for noninvasive
subsurface inspections of both structural and biological media. Current
ultrasound probes are rigid and bulky and cannot readily image through nonplanar
three-dimensional (3D) surfaces. However, imaging through these complicated
surfaces is vital because stress concentrations at geometrical discontinuities
render these surfaces highly prone to defects. This study reports a stretchable
ultrasound probe that can conform to and detect nonplanar complex surfaces. The
probe consists of a 10 × 10 array of piezoelectric transducers that exploit an
“island-bridge” layout with multilayer electrodes, encapsulated by thin and
compliant silicone elastomers. The stretchable probe shows excellent
electromechanical coupling, minimal cross-talk, and more than 50%
stretchability. Its performance is demonstrated by reconstructing defects in 3D
space with high spatial resolution through flat, concave, and convex surfaces.
The results hold great implications for applications of ultrasound that require
imaging through complex surfaces.
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INTRODUCTION
Ultrasound imaging technologies have been widely used to visualize internal
discontinuities in objects for nondestructive evaluation, structural health
monitoring, and medical diagnosis because of their noninvasiveness, high
accuracy, great sensitivity, and strong penetration capabilities (1–3).
Ultrasound probes with flat bases have been created to successfully accommodate
different components of planar surfaces. However, these rigid probes cannot
achieve a solid interfacial contact and therefore good coupling with irregular
nonplanar surfaces, which are ubiquitous in real objects. Air gaps at these
interfaces lead to large acoustic energy reflections and wave distortions,
thereby creating unreliable testing results (fig. S1) (4). Ultrasonic couplants,
such as water and gel, are typically used to remove the air gaps. However, an
abundant use of the couplants will lead to a high-pass filter effect of the
ultrasonic signals, causing significant canceling of small response echoes (5).
Furthermore, extensive use of the couplants will bring about an ~80% incident
energy transmission loss at the interface between the couplant and the subject
because of the significant mismatch of their acoustic impedances (6). In
addition, these rigid and bulky probes cannot be applied to hard-to-reach
locations such as small cavities and slits. Thus, components at these locations
normally have to be disassembled for a reliable diagnosis. At the same time, the
stress concentrations present at the geometrical discontinuities of load-bearing
objects make these regions particularly prone to defects (7). Although many
methods have been reported to solve this interfacial coupling problem (8, 9), a
number of disadvantages of the existing approaches remain, such as limited
specimen size (10), demanding probe offset (11), and bulky probe housing (12),
all of which compromise the feasibility of in situ detection, detection accuracy
and sensitivity, and operation convenience of ultrasonic measurements.
Recent efforts have focused on developing flexible ultrasonic probes that can be
divided into three main categories: using organic piezoelectric films as
transducers, embedding piezoelectric ceramic into polymer substrates, and
fabricating capacitive micromachined ultrasonic transducers (CMUTs). The organic
piezoelectric films have good flexibility. However, the polymer piezoelectrets,
typically polyvinylidene fluoride and its copolymer films (13), are not suitable
to serve as transmitters because of their low electromechanical coupling
coefficients (a parameter that characterizes the coupling between electrical
energy and mechanical energy), low dielectric constants, and high dielectric
losses (14). Moreover, their low Curie points make them difficult to process,
and high-temperature applications result in phase transformations, which
completely degrade the piezoelectric properties (15). The piezoelectric ceramics
produce superior electromechanical performance and ease of processing. However,
they cannot conform to curved surfaces without external forces because of the
large elastic moduli of substrates (4, 15–17). The external force, usually
applied manually, is often inconsistent. As a result, noise or even artifacts in
the acquired pulse-echo signals can arise because of variations of the coupling
conditions at the transducer-specimen interface. Moreover, for some applications
related to long-term structural condition monitoring, such as fatigue crack
growth at hidden or hard-to-access places of aircrafts and steamboats, the
mechanical robot cannot support the testing (18). The CMUTs are fabricated on
disjoined silicon wafers, and polydimethylsiloxane (PDMS) refilling the trenches
among the elements makes the transducers flexible (17). This passive polymer
filler compromises their conformability on curved surfaces. Besides, the silicon
substrates are likely to be secondary resonators that generate longitudinal
waves with unwanted frequencies and eventually result in artifacts in the images
(19). Also, CMUTs generally have a lower electromechanical efficiency than
piezoelectric ceramics due to inhomogeneity and parasitic capacitances among the
arrayed elements (20, 21). In all cases, these flexible probes can only conform
to developable surfaces (such as cylindrical surfaces), not to nondevelopable
surfaces (such as spherical surfaces). In addition, the flexible conductive
interconnections are subject to breaking or debonding when repeatedly used (22)
because being flexible is insufficient to accommodate the sophisticated and
time-dynamic motion of the electrodes and the device during the measurements.
These drawbacks represent a bottleneck for the development of advanced probes
that combine excellent ultrasonic performance with desirable mechanical
properties that allow for application to general complex surfaces.
Here, we report a low-profile membrane-based stretchable ultrasonic probe. The
probe exploits an array of thin and high-performance 1-3 piezoelectric
composites as transducers, multilayered serpentine metal traces as electrical
interconnects, and low-modulus elastomer membranes as encapsulation materials.
The resulting device has a high electromechanical coupling coefficient (keff ~
0.60), a high signal-to-noise ratio (SNR; ~20.28 dB), a wide bandwidth
(~47.11%), a negligible cross-talk level between adjacent elements (~−70 dB),
and a high spatial resolution (~610 μm) at different depths. The “island-bridge”
layout offers biaxial stretchability of more than 50% with minimal impact on the
transducer performance, which allows the device to work on nonplanar complex
surfaces. With these unique properties, the device can obtain three-dimensional
(3D) images of complex defects under flat, concave, and convex surfaces.
RESULTS
DESIGN AND CHARACTERIZATION OF THE STRETCHABLE ULTRASONIC ARRAYS
The schematic device structure is shown in Fig. 1A. The piezoelectric
transducers are arranged in a 10 × 10 array, connected by an island-bridge
structured matrix (the schematic fabrication processes are shown in fig. S2).
Each island hosts a rigid element. The wavy bridges can unfold to accommodate
the externally applied strain, with limited strain in the components themselves.
Therefore, the matrix is rigid locally but soft globally. Each element in the
array is individually addressable. The soft probe can consequently reconstruct
the target morphology in multisection images. Figure 1B shows the exploded view
of one element. Both the substrate and superstrate are silicone elastomer thin
films, whose low modulus (~70 kPa) and large stretchability (~900%) offer an
extremely compliant platform to accommodate a diverse class of building blocks,
such as piezoelectric elements, metal interconnects, backing layers, and solder
paste. The thickness of the elastomer substrate and superstrate is 15 μm to
provide both high acoustic performance (23) and mechanical robustness of the
device (figs. S3 and S4). The islands and bridges are patterned bilayers of Cu
(20 μm)/polyimide (PI; 2 μm). The PI layer greatly enhances the bonding strength
between the Cu and the elastomer.
Fig. 1 Schematics and design of the stretchable ultrasonic transducer array.
(A) Schematics showing the device structure. (B) Exploded view to illustrate
each component in an element. (C) The optical image (bottom view) of four
elements, showing the morphology of the piezoelectric material and bottom
electrodes. (D) The tilted scanning electron microscopy image of a 1-3
piezoelectric composite. (E) The optical image (top view) of four elements,
showing the morphology of the backing layer and top electrodes. (F to H) Optical
images of this stretchable device when (F) bent around a developable surface,
(G) wrapped on a nondevelopable surface, and (H) in a mixed mode of folding,
stretching, and twisting, showing its mechanical robustness.
Open in viewer
Piezoelectric 1-3 composites are chosen as the active material of the
transducers (Fig. 1, C and D). Compared with an isotropic lead zirconate
titanate (PZT), the anisotropic 1-3 composites have superior electromechanical
coupling coefficients (thickness mode) that convert the majority of electrical
energy to vibrational energy. In addition, the surrounding epoxy filler
effectively suppresses transverse vibrations of PZT pillars (fig. S5) (24),
leading to enhanced longitudinal waves that go into the targeted objects. Molded
blocks of a Ag-epoxy composite serve as the backing layer (Fig. 1E). The backing
layer effectively dampens the ringing effects (excessive vibrations) of the
piezoelectrics, which shortens the spatial pulse length and broadens the
bandwidth (fig. S6) (25) and thus improves the image axial resolution. Ag epoxy
and solder paste are used to build robust and electrically conductive 1-3
composite/backing layer and 1-3 composite/metal electrode interfaces,
respectively. Because of the close acoustic impedances of the 1-3 composite [~20
megarayleigh (MR)] and the targets to be tested (Al, ~18 MR), the matching layer
is not necessary in this study (26, 27).
On one hand, the pitch between adjacent transducer elements should be small to
reduce side lobe and grating lobe artifacts in the acquired images (28). On the
other hand, sufficient space between the elements should be allocated to the
serpentine interconnects for sufficient stretchability. A pitch of 2.0 mm (1.2
mm × 1.2 mm element footprint with a spacing of 0.8 mm) can achieve more than
30% reversible stretchability. The high spatial resolution (~610 μm), negligible
cross-talk level between adjacent elements (~−70 dB), and artifact-free images
validate this pitch design. Within these limited footprints, the island-bridge
electrode layout design is critical, considering the large number of electrical
connections needed for wiring the 10 × 10 array. An active multiplexing matrix
under the ultrasound transducers could be a potential solution (29). However,
the structural support materials introduced by the multiplexing matrix will
negatively affect device stretchability. Multilayered electrodes have been
demonstrated (30, 31), but the electrode design, passive dielectrics, and the
substrate only makes the devices flexible, not stretchable. To individually
address the 100 transducer elements, a minimum of 101 electrodes with a common
ground electrode is needed. It is very challenging to place this large number of
electrodes within limited footprints using conventional single-layer designs.
Thus, we invented a multilayered electrode design based on the “transfer
printing” method, which greatly enhanced the level of device integration
compared to single-layer designs. This design consists of five layers of
“horseshoe”-configured serpentine electrodes. One electrode lies at the bottom
of the transducers as the common ground layer (fig. S7). The other 100
electrodes are well aligned and distributed into four layers on top of the
transducers as stimulating electrodes (fig. S8). Thin films of silicone
elastomer (35 μm thick) provide insulation and adhesion between adjacent layers.
The central area of each layer is selectively protected using customized masks
during fabrication to allow the islands (bonding pads) to be exposed to the
array elements (fig. S9). Laser ablation is used to quickly pattern serpentine
structures (figs. S10 and S11). This method has been mostly focused on rigid or
flexible substrates, but has been seldom focused on silicone substrates for
stretchable electronics. The challenges for using on stretchable substrates are
(i) controlling the laser power to fully ablate the pattern while avoiding the
pattern delamination from the temporary PDMS substrate and (ii) tuning the
surface tackiness of the temporary PDMS substrate to allow the subsequent
transfer printing of the patterned electrodes. We solved these challenges and
developed a fabrication protocol for stretchable electronics using laser
ablation (see Materials and Methods for details). Compared with microfabrication
methods by lithography and etching (32, 33), which require sophisticated
fabrication processes, chemicals, shadow masks, and a clean room environment,
laser ablation is time-efficient, low-cost, and offers high throughput. The
as-fabricated final device is seen in Fig. 1 (F to H), which highlights its
excellent mechanical properties when conforming to developable (cylindrical) and
nondevelopable (spherical) surfaces, and under mixed modes of folding,
stretching, and twisting. The device can easily achieve conformal contact with
various nonplanar surfaces of real components, such as pipeline elbows, wheel
edges, and rail tracks (fig. S12). An anisotropic conductive film (ACF) bonded
to the Cu interconnects offers conductive access to external power supplies and
data acquisition systems (fig. S13). Details of materials and fabrication
processes are provided in Materials and Methods.
ELECTROMECHANICAL CHARACTERIZATIONS
Ultrasound emission and sensing rely on the reversible conversion of mechanical
and electrical energy. The electromechanical coupling capability is thus a key
metric to evaluate the ultrasound transducer performance. As illustrated in Fig.
2A, the electrical impedance and phase angle spectra of the 1-3 composite before
and after fabrication are measured, from which we can obtain the
electromechanical coupling coefficient k (kt and keff) and the degree of poling,
respectively (34). The black curves show two sets of well-defined peaks,
corresponding to the resonance frequency fr and the antiresonance frequency fa.
Accordingly, the kt and keff of the 1-3 composite before and after the
fabrication are calculated to be ~0.55 and ~0.60, respectively (see Materials
and Methods for details). The phase angle of the 1-3 composite at the central
frequency slightly dropped from ~60° before fabrication to ~50° after
fabrication because of the heat-induced slight depolarization of the 1-3
composite. The final phase angle of ~50°, which significantly exceeds many
previous reports in flexible or rigid ultrasound probes, due to the intrinsic
properties of the 1-3 composite material and optimized fabrication processes
(16, 35), demonstrates that most of the dipoles in the 1-3 composite align
during poling, thereby indicating the outstanding electromechanical coupling
properties of our device.
Fig. 2 Characterizations of piezoelectric and mechanical properties.
(A) The impedance and phase angle spectra of the 1-3 composite before and after
processing, showing good electromechanical coupling of the fabricated transducer
(keff, ~0.60; θ, ~50°). (B) Pulse-echo response and frequency spectra, with a
short spatial pulse length (~1.94 μs), a high SNR (~20.24 dB), and a wide
bandwidth (~47.11%). (C) The resonance and antiresonance frequency variations of
the 100 transducer elements. The mean values/SDs are 3.51 MHz/56.8 kHz
(resonant) and 4.30 MHz/59.1 kHz (antiresonant), respectively. The 100% yield
demonstrates fabrication robustness. (D) Average cross-talk levels between
elements that are adjacent, two elements away, and three elements away, showing
the outstanding anti-interference capacity of the device. (E) The optical image
(left) and corresponding finite element analysis (FEA) simulation (right) of a 2
× 2 array under 50% biaxial tensile strain, showing its excellent
stretchability. The local strain level (maximum principal strain) in the
interconnects is indicated by the color scale. (F) The optical image after
releasing the biaxial strain of 50%. The zoomed-in image highlights plastic
deformation and local delamination of the interconnects upon loading/unloading.
(G) Electrical impedances of the transducer under different strain levels,
showing the mechanical stability of the device.
Expand for more
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The Krimholtz-Leedom-Matthaei (KLM) model (36, 37) in MATLAB allows for the
prediction of the impulse response of the transducer (see section S1), as a
theoretical validation for our device design. The simulated results demonstrate
the superb performance of the device in terms of spatial pulse length,
bandwidth, and SNR (fig. S14). Figure 2B shows the experimental results of
pulse-echo response and its frequency spectrum. The pulse-echo response, with a
narrow spatial pulse length (~1.94 μs), a wide bandwidth (~47.11%; see Materials
and Methods for details), and a high SNR (~20.24 dB), matches well with the
simulation result (fig. S14) and is on par with that of commercial flexible
ultrasonic transducers (38). The outstanding transducer performance results from
(i) the excellent electromechanical coupling of the transducer and (ii) the
optimized backing layer that reduces ringing effects.
The impedance measurements enable extraction of the resonant and antiresonant
frequencies of each element in the 10 × 10 array (Fig. 2C). All 100 elements are
functional. The mean values are 3.51 MHz (resonant) and 4.30 MHz (antiresonant),
with SDs of 56.8 and 59.1 kHz, respectively. The stable capacitance (~37.28 pF)
and low dielectric loss (tan δ < 0.02) of the array (figs. S15) further suggest
a remarkable uniformity across the array and a reliable fabrication method (fig.
S16). Another important metric that assesses the performance of the array is the
cross-talk, which indicates the degree of interference between the elements.
Figure 2D shows the cross-talk between elements with different spacings. All
cross-talk levels are around −70 dB, with slight fluctuations, which is
significantly lower than the standard −30 dB in the field (39). The outstanding
anti-interference properties arise from the 1-3 composites’ effective
suppression of spurious shear and from the silicone elastomer providing
effective isolation among the elements. Overall, this combination of properties
ensures low levels of noise in the ultrasonic imaging system (25).
MECHANICAL CHARACTERIZATIONS
Mechanical properties of conductive interconnects are critical for flexible and
stretchable devices (40–42). Experimental results from biaxial stretching of the
layered structures of serpentines between 0 and 50% and corresponding 3D finite
element analysis (FEA) are shown in Fig. 2E and fig. S17. A 2 × 2 array of
elements is selected for visualization of the key mechanics involved (see
Materials and Methods for details). Under tensile loading, the horseshoe
serpentines undergo an in-plane unraveling process and out-of-plane rotation and
twisting, both of which mitigate the level of strain in the islands themselves
(43, 44). Specifically, in these ultrasound arrays, 50% biaxial stretching
produces a maximum of only ~1.2% tensile strain in the Cu interconnects, as
shown in the FEA image of Fig. 2E. After the serpentines have fully unraveled
(that is, finished rotating in-plane), the tensile strain in the Cu
interconnects increases rapidly, thus defining the stretching limit of the
serpentines (43, 45), which is between ~50 and 60% in this case. Going beyond
this limit will lead to fracture of the serpentines. In addition, for the
reliability of these devices, they must be capable of sustaining mechanical
integrity upon repetitive loading. In metals such as Cu, cycling into the
plastic regime will cause permanent deformation of the interconnects, which may
affect device performance or may eventually produce fatigue cracks (46).
According to both the simulations and the experiments, ~30 to 40% biaxial
stretching produces irreversible deformation in the serpentines upon releasing
(fig. S17) and partial delamination between the serpentines and the silicone
elastomer, as highlighted in Fig. 2F. However, below 30 to 40% biaxial
stretching, mechanical integrity is maintained. Moreover, mechanical
deformations have minimal influence on device performance, which is reflected by
the stable impedances of each element (Fig. 2G and figs. S18 and S19) and the
resistance of the serpentines (fig. S20) at various levels of tensile strains
and bending curvatures.
SPATIAL RESOLUTION CHARACTERIZATION
One of the important performance metrics of ultrasound imaging systems is the
spatial resolution, in both axial and lateral directions. For the stretchable
ultrasound probe, the axial resolution remains constant under different bending
curvatures at a defined resonant frequency and bandwidth of the transducer. The
lateral resolution is mainly dependent on device geometry, which affects the
focal length and aperture size. The f number is used to define the ratio between
the focal length and the aperture size (47). To comprehensively explore the
lateral resolution of the probe with various f numbers, we performed a series of
imaging experiments in which the ultrasonic probe is bent to different
curvatures. As shown in Fig. 3A, the spatial resolution was evaluated by
focusing the array at focal lengths of 20, 32, 37, and 52 mm, respectively, to
image a Cu wire (300 μm in diameter) located at a particular focal point in a
phantom sample. The image was reconstructed using the DMAS (delay multiply and
sum) algorithm (48), which more effectively suppressed the level of noise floor
(~−40 dB, causing the energy ratio of noise to reflector to be only 0.01%)
compared with a conventional algorithm such as DAS (delay and sum; Fig. 3B)
(49). In light of this metric, the side lobes and grating lobes in images can be
greatly reduced by using the DMAS, and the results from these four tests are
combined in Fig. 3C. A configuration of −10 dB dynamic range in combination with
an image resolution of 20 pixels/mm is applied to highlight the imaging
capabilities. The imaging principle of the DMAS algorithm and its detailed
comparison with DAS are elaborated on in Materials and Methods and section S2.
Fig. 3 Characterization of spatial resolution.
(A) Schematics of spatial resolution measurement setup, with focal lengths of
20, 32, 37, and 52 mm, respectively. (B) Comparison of noise floors
reconstructed by DMAS and DAS algorithms, revealing the benefits of the DMAS
algorithm with only 0.01% energy ratio of noise to the reflector. (C) Images of
wire phantom combining the four tests with different f numbers, showing the
capability of focusing at different depths and obtaining high-resolution images.
(D) Axial and (E) lateral line spread functions for the center wire at different
focal lengths. Resolution is defined as the linespread function width at an
intensity of −6 dB. (F) Experimental (Exp.) and simulation (Simu.) results of
lateral and axial resolutions.
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Plots of the axial and lateral line spread functions of the obtained images
(Fig. 3B) are shown in Fig. 3 (D and E). The measured FWHM (full width at half
maximum) resolutions (−6 dB) (47, 50) were calculated for the axial and lateral
directions, as indicated by the dashed lines. As the f number decreases, the
axial resolution remains relatively constant at around 610 μm, and the lateral
resolution improves approximately linearly from 789 to 344 μm. These results are
in line with the theoretical results (an axial resolution of around 601 μm;
lateral resolutions ranging from 787 to 284 μm) from the MATLAB k-Wave toolbox
simulations (Fig. 3F). The fine spatial resolution at the focal point, which is
comparable to the 3.5-MHz commercial ultrasound probe resolution of 610 μm (51),
is due to the combined effects of the high-performance transducers, a strategic
device structural design, and an advanced imaging algorithm.
MULTIVIEW IMAGING ON COMPLEX SURFACES
We used the stretchable ultrasonic device to image customized Al work pieces
with embedded defects under planar, concave, and convex surfaces. The detailed
experimental setup and method are discussed in the Materials and Methods and in
figs. S21 and S22. In all cases, a straight defect (2 mm in diameter, orthogonal
to the side surface) was created with different distances from the top surface
(Fig. 4, A to C, first column). The device was laminated seamlessly on the test
surfaces. The synthetic aperture focus (SAF) method was applied to reconstruct
the corresponding images (section S2) (50). This method allows a sparse
transmitter-receiver scheme that bypasses the need for simultaneous excitations,
minimizing the number of simultaneously active elements while preserving the
image quality. As indicated by the wave field simulation results, the main lobes
of the transducer are parallel, divergent, and focused for the planar, concave,
and convex surfaces, respectively (Fig. 4, A to C, second column, and movies S1
to S3). Considering that the central defect acts as a secondary wave source and
the transducer is primarily sensitive to out-of-plane motion (direction normal
to transducer’s sensing surface), the target surface curvature can greatly
influence the captured signal strength. Specifically, for the convex surface,
most of the reflected longitudinal wave motion from the defect aligns with the
direction perpendicular to the sensing surface; for the concave surface, the
reflected wave motion aligns with the in-plane motion (direction parallel to the
sensing surface); for the planar surface, which is an intermediate case, the
sensitivity of the transducer mainly depends on the component of the reflected
wave vector normal to the sensing surface. To acquire the defect signals, for
each case, we obtained 90 sets of data. Longitudinal wave reflection signals
from the defects and the bottom boundaries, with more than 18 dB SNR, can be
collected with predicted times of arrival (Fig. 4, A to C, third column). The
obtained full-field images of the defects are shown in the fourth column of Fig.
4 (A to C), which have no artifacts and match the simulation results extremely
well (fig. S23). These results suggest that the stretchable ultrasound probe is
capable of accurately imaging defects in media of complex surface geometries.
Fig. 4 Two-dimensional images of a linear defect under complex surfaces.
Optical images of experimental setups with the stretchable ultrasonic device
tested on (A) planar, (B) concave, and (C) convex surfaces showing the good
conformability of the device on these surfaces (first column); simulation
results showing the different wave fields and sensing modes (second column);
pulse-echo signals from the defects and boundaries with high SNR (third column);
and acquired 2D images using DMAS algorithms with accurate and artifact-free
positions (fourth column). S wave, shear wave; L wave, longitudinal wave; SNRD,
SNR of pulse-echo response from the defect.
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For practical engineering inspections, detection of multiple defects is of
particular interest, for example, welding inspection of a pipeline (52) and rail
track detection under shelling (53). The stretchable ultrasonic device was used
for 3D internal structure visualization by imaging two defects with different
depths and orientations under a sinusoidal curved surface. A schematic of the
experimental setup is shown in Fig. 5A, with one defect orthogonal to the x-z
plane at a depth of 4.0 cm below the top surface, and the other defect 18°
tilted away from the y axis at a depth of 6.0 cm below the top surface. Each 1 ×
10 linear array in the x-z plane generates a 2D cross-sectional image of the two
defects using the DMAS algorithm (fig. S24), similar to Fig. 4. The upper defect
reflects part of the wave and reduces the wave energy reaching the lower defect.
Thus, it produces a shadowing effect (54), which is exacerbated by the tilted
configuration of the lower defect as the array scans from the y = 0 to the y =
1.8 plane. The 3D image can be reconstructed by integrating the 10 slices with a
2-mm pitch along the y axis, as shown in Fig. 5B. The shadowing effect is
removed by normalizing against the peak intensity of each defect. The
corresponding front, top, and side views are shown in Fig. 5 (C to E), which
accurately match the design in Fig. 5A, thereby demonstrating the capability of
volumetric imaging using the stretchable ultrasonic probe. Similar protocols of
testing and imaging reconstruction can be applied to general and more
sophisticated surfaces.
Fig. 5 Three-dimensional image reconstruction of intricate defects under a
convex surface.
(A) Schematics of the experimental setup, illustrating the spatial location and
relative orientation of the two defects in the test subject. (B) The
reconstructed 3D image, showing complete geometries of the two defects. (C to E)
The 3D image from different view angles, showing the relative positions and
orientations of the two defects to the top surface, which match the design well.
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DISCUSSION
The hybridized material integration, electrode design strategies, and imaging
algorithm development introduced here provide a foundational basis for
stretchable ultrasound imaging arrays that allow nondestructive 3D volumetric
inspections under general complex surfaces. The high-performance anisotropic 1-3
piezoelectric composites suppress shear vibrations, reduce cross-talk among the
transducer, enhance longitudinal vibrations, and thus improve the overall
sensitivity and SNR. Five-layered serpentine electrodes enable a high level of
integration and large stretchability of more than 50%. The stretchable
ultrasound probe, consisting of a 10 × 10 array of individually addressable
transducer elements, can focus at different depths, with comparable spatial
resolutions with existing rigid probes. The unique device design, combined with
the advanced DMAS imaging algorithms, enables accurate, artifact-free,
full-field, and nondestructive examinations underneath general complex surfaces.
Future studies will focus on improving the device performance by reducing the
pitch between elements and exploring the integration of more active elements,
incorporating matching layers, automated sensor positioning systems, and
wireless signal actuation and transmission chips, for distributed, mobile, and
real-time subsurface health monitoring of infrastructures and the human body.
MATERIALS AND METHODS
FABRICATION OF THE FIVE-LAYER ELECTRODES
The process began with a coating of PI (2 μm thick) on Cu sheets (20 μm thick).
PI [poly(pyromellitic dianhydride-co-4,40-oxydianiline) amic acid solution,
PI2545 precursor, HD MicroSystems] was first spin-coated on the Cu sheets
(Oak-Mitsui Inc.) at 4000 rpm for 60 s (MicroNano Tools). Then, PI/Cu was baked
on a hotplate at 110°C for 3 min and 150°C for 1 min sequentially and then fully
cured in a nitrogen oven at 300°C for 1 hour. A glass slide coated with a layer
of PDMS (20:1, Sylgard 184 silicone elastomer) served as a substrate for
laminating the PI/Cu sheet. The PI and PDMS were activated for bonding by
ultraviolet (UV) light (PSD series Digital UV Ozone System, Novascan) for 1.5
min. Five separate pieces of Cu sheets were then patterned in
island-bridge–structured geometries (designed by AutoCAD software) by pulsed
laser ablation (Laser Mark’s). The laser parameters (central wavelength, 1059 to
1065 nm; power, 0.228 mJ; frequency, 35 kHz; speed, 300 mm/s; and pulse width,
500 ns) were optimized to process Cu with the highest yield. Thin silicone
superstrates/substrates of devices (15 μm each; Ecoflex-0030, Smooth-On) were
prepared by mixing two precursor components together in a 1:1 ratio,
spin-coating at 4000 rpm for 60 s, and curing at room temperature for 2 hours.
Here, PDMS was used as a temporary substrate where the PI/Cu sheet was laminated
for laser ablation. Compared with the PDMS, Ecoflex has a lower Young’s modulus
[Young’s moduli of Ecoflex-0030 (1:1) and PDMS (20:1) are ~60 kPa (55) and ~1
MPa (56), respectively]. Thus, we chose Ecoflex as the substrate, superstrate,
and filler in our device to ensure the low modulus of the device that allowed
intimately conforming to the highly curved surfaces.
For the first layer, water-soluble tape (3M) was used to transfer-print the
patterned Cu electrode to the Ecoflex superstrate after 3 min of UV activation
(57). A separate piece of water-soluble tape was used to selectively mask the
connect pads at the center and top of the electrode that would be exposed to
bond the transducer array and ACF cables (Elform). Next, a 35-μm-thick Ecoflex
film was spin-coated at 3000 rpm for 60 s and cured at 80°C for 20 min, forming
an insulating layer while the Ecoflex on top of the water-soluble tape mask was
removed by dissolving the water-soluble tape. Subsequent layers of electrodes
were laminated, with alignment to the previous layer of electrodes, in a similar
manner. The integrated four-layer top electrodes are shown in fig. S8. The
bottom electrode was fabricated and transfer-printed to a separate Ecoflex
substrate (fig. S7). Finally, ACF cables were hot-pressed onto the electrodes to
serve as the connection access for data communication and power supply (fig.
S13).
ASSEMBLING OF TRANSDUCER ARRAYS AND THEIR INTEGRATION WITH ELECTRODES
As shown in fig. S2, the process began with the fabrication of the backing layer
and the 1-3 composite (Smart Material Corp.). The conductive backing layer was
prepared by mixing a Ag-epoxy composite with a hardener (E-Solder 3022, Von
Roll) in a 12.5:1 ratio and then curing the mixture at 60°C for 8 hours. The
backing layer thickness was fixed at 580 μm by mounting the backing layer
precursor between two glass slides. The backing layer was then diced into pieces
of 1.2 mm × 1.2 mm by a dicing saw (DAD3220, DISCO). The 1-3 composites were
fabricated from PZT ceramics and epoxy using the dice-and-fill technique. The
size of each PZT pillar was 100 μm × 100 μm with a spacing of 55 μm (Fig. 1D).
Each one of the 1-3 composite elements was diced to 1.2 mm × 1.2 mm and bonded
with the backing layer via Ag-epoxy (EPO-TEK H20E, Epoxy Technology) at 150°C
for 5 min. The single-layer bottom electrode was bonded with the 10 × 10 arrayed
1-3 composite, using a customized scaffold, by solder pastes [Sn42Bi57.6Ag0.4
(melting point, 138°C), Chip Quik Inc.] cured in the oven at 150°C for 6 min.
The same approach was used to bond the four-layer top electrode with the backing
layer. The gap between the sandwiched device was then filled by Ecoflex and
cured at room temperature for 2 hours. Afterward, the glass slides were removed,
yielding a free-standing stretchable ultrasound transducer array.
ELECTROMECHANICAL AND MECHANICAL TESTING OF THE DEVICE
A high-voltage power supply (Model 355, Bertan), with a direct voltage output of
52.38 kV/cm, provided a platform to polarize the device for 15 min. The
polarization hysteresis loop (fig. S25) was measured to determine the minimal
voltage needed to fully polarize the 1-3 composite in the silicone medium
without electrical breakdown. A network analyzer (Agilent Technologies) with a
scanning range of 2 to 6 MHz under the Smith mode gave the impedance and phase
angles of the transducer. Electromechanical efficiency is a parameter that
characterizes the degree of energy coupling efficiency between the electrical
and mechanical forms. The electromechanical coupling coefficient k is the factor
that quantitatively evaluates electromechanical efficiency. The
electromechanical coupling coefficients of the 1-3 composite and the transducer
(the transducer contains the 1-3 composite and the backing layer, and is
processed by heating and poling), kt and keff, were derived from Eqs. 1 and 2,
respectively (58)
kt=π2frfa tan(π2fa−frfa)−−−−−−−−−−−−−−−−−−√kt=π2frfa tan(π2fa−frfa)
(1)
keff=1−f2rf2a−−−−−−√keff=1−fr2fa2
(2)
where the resonant frequency fr and the antiresonant frequency fa were extracted
from the impedance and phase angle spectra. An experimental system, including a
pulser-receiver (model Panametric 5077PR, Olympus), an oscilloscope (LeCroy
WaveJet 314), and a 300-μm-diameter copper wire in the phantom, was used to
obtain the pulse-echo signal and frequency spectra. The frequency bandwidth (BW)
of the signal at −6 dB was determined by Eq. 3
BW =fu−flfc×100%BW =fu−flfc×100%
(3)
where fu is the upper frequency, fl is the lower frequency, and fc is the
central frequency (47). A function generator (AFG3521, Tektronix) and an
oscilloscope (LC534, LeCroy Corp.) were used to assess the cross-talk. The
sinusoid burst mode with a peak-to-peak voltage of 5 V was used to excite the
elements in the array. The frequency was scanned between 2 and 6 MHz with a step
size of 0.2 MHz. The cross-talk level was then defined by counting the ratio of
the peak voltages to the reference voltage (the voltage under a 1-megohm
coupling on the oscilloscope). The capacitance and dielectric loss of array
elements were measured by an inductance-capacitance-resistance (LCR) digital
bridge machine (QuadTech).
Mechanical testing of a 2 × 2 transducer array was performed with a customized
biaxial stretcher. To accurately evaluate the biaxial stretchability, the strain
was quantified based on the distance between the two electrodes. Images of the
device under different strain levels were collected with a charge-coupled device
(OMAX) on an optical microscope (AmScope). The electric impedance of the
transducer under stretching and bending states and the relative resistance
change of the Cu serpentines at various tensile strain levels were tested by a
network analyzer and a source meter (Keysight Technologies), respectively.
FEA SIMULATIONS
The commercial software package ABAQUS allowed the simulation of the mechanical
response of the transducer arrays. The composite layer (Ecoflex, Cu, and PI)
consisted of hybrid hexahedral elements (C3D8H). The simulations used values of
the elastic moduli of Ecoflex, PI, and Cu of 0.06, 2300, and 41,500 MPa,
respectively. To obtain the yield strength value of the Cu, four pieces of Cu
slides with a high aspect ratio (width, 4.18 mm; thickness, 0.02 mm; length,
18.78 mm) were measured by tensile testing. The testing rate was 1% tensile
strain per minute, and the load cell was 1 kN (Instron 5965). The stress-strain
curves were obtained where the yield strength value of Cu slides (187 MPa) was
extracted and used in the simulations.
NONINVASIVE INSPECTION OF THE INTERNAL DEFECTS
To test specimens with planar, concave, and convex surfaces, the coordinate
location of each transducer element in the array was determined by aligning the
device with a known marker position on the surface. For more complex surfaces, a
3D scanner can be used to locate the exact coordinates of each element. For the
specimens whose surfaces were sinusoidal, the amplitude was 4 mm peak to peak,
and the wavelength was 40 mm. A data acquisition system, composed of a
pulser-receiver, an oscilloscope, a switch controller (NI USB-6008), and
high-voltage chips (HV2601), was developed using LabVIEW (National Instruments),
as shown in figs. S21 and S22. We used LabVIEW to program the switch controller
so that it could turn on and turn off the high-voltage chip. This system allowed
the device to automatically transmit and receive ultrasound signals. An
electrical impedance matching system between the transducer and the
pulser-receiver (50 ohms) was used to minimize power reflection when exciting
the transducer. The enhanced power transmission efficiency improved the SNR
(fig. S26). A step-pulse excitation (−100 V) was applied on the selected
element, and the responses from the other elements were collected. By iterating
this procedure over a 1 × 10 linear array, 90 sets of waveform data could be
acquired. The received raw signals were first filtered with a band-pass filter
based on the continuous wavelet transform to remove the unwanted frequency
components and minimize the filter-induced phase shift (fig. S27). Each filtered
waveform was decomposed into its in-phase and phase-quadrature components
through the Hilbert transform, and an improved SAF algorithm was then applied to
each of the Hilbert transformed components separately. Finally, the pixel
intensity of the final images was reconstructed by computing the moduli of the
two components from the Hilbert transform and converting it to the decibel
dynamic range. The image resolution of 20 pixels/mm, which fulfills the
half-wavelength spatial sampling of longitudinal waves in the phantom at 3.5 MHz
(Nyquist-Shannon sampling theorem), was applied here to completely preserve the
information in the signals.
Here, the SAF technique based on DMAS was implemented for image reconstruction
(48). To reconstruct an image I(x, y) at each pixel P(x, y) with DMAS,
considering a linear array of 1 × M elements, M − 1 ultrasound signals were
recorded each time when one element was activated as the transmitter and the
remaining M − 1 elements were the receivers. Thus, a total of M ∙ (M − 1)
signals were obtained. The amplitudes of the received signals (A) were
appropriately backpropagated for each combination of transmitter and receiver.
Once all the signals were in phase with regard to pixel P(x, y), they were
combinatorially coupled and multiplied. If the number of received signals was N,
then the number of multiplications to be performed was given by all the possible
signal pair combinations (N2)=N2−N2(N2)=N2−N2. The backpropagated DMAS algorithm
can be written as (48)
IDMAS(x,y)=∑N−1i=1∑Nj=i+1sign[Ai(τi,xy)Aj(τj,xy)]*|Ai(τi,xy)Aj(τj,xy)|−−−−−−−−−−−−−−√IDMAS(x,y)=∑i=1N−1∑j=i+1Nsign[Ai(τi,xy)Aj(τj,xy)]*|Ai(τi,xy)Aj(τj,xy)|
(4)
where Ai and Aj are the signals received by the ith and jth transmitter-receiver
pairs, respectively, and τi,xy and τj,xy are the backpropagation time
corresponding to the travel time of the wave from the ith and jth
transmitter-receiver pairs, respectively, through the focus point P(x, y). DMAS
suppressed the level of noise floor to −40 dB. Thus, the energy ratio of noise
to the reflector was 0.01%, which can be calculated by
−40=10*lgP1P0−40=10*lgP1P0
(5)
where P1 and P0 are noise and reflector energies, respectively.
ACKNOWLEDGMENTS
We thank S. Cai’s group at the University of California San Diego (UCSD) for the
Young’s modulus testing, Y. Dai for analyzing the strain distribution of the
device in stretching states, and S. Xiang for constructive feedback on the
manuscript preparation. Funding: The project described was partially supported
by the NIH (grant R21EB025521) and Clinical and Translational Science Awards
funding (grant UL1TR001442). Additional partial support was provided by the UCSD
Center for Healthy Aging, the U.S. Federal Railroad Administration through grant
FR-RRD-0027-11, and the U.S. NSF through grant CMMI-1362144. The content is
solely the responsibility of the authors and does not necessarily represent the
official views of the NIH or NSF. Author contributions: H.H., X.Z., Chonghe
Wang, L.Z., X.L., F.L.d.S., and S.X. designed and executed the experiments,
reconstructed images, wrote the paper, and were responsible for Figs. 2 (A, B,
F, and G), 3, 4, and 5. S.L. and M.P. carried out FEA simulations and were
responsible for Fig. 2E. R.C., Z.C., and Q.Z. characterized the
electromechanical coupling performances and were responsible for Fig. 2 (C and
D). Z.H., Chunfeng Wang, and Y. Gu edited the manuscript, characterized
mechanical properties, and were responsible for Fig. 1 (F to H). Y.C., Y. Lei,
and T.Z. were responsible for Fig. 1 (C to E). Chonghe Wang and N.K. were
responsible for Fig. 1 (A and B). Y. Guo, W.Z., Y. Li, and A.N. made
contributions in device design and fabrication. Y.T. and S.S. made contributions
in algorithm development. Competing interests: S.X., L.Z., Chonghe Wang, H.H.,
and X.L. are inventors on a Patent Cooperation Treaty (PCT) patent application
related to this work filed by the Regents of the University of California
(PCT/IS18/13116, filed 10 January 2018). The other authors declare that they
have no competing interests. Data and materials availability: All data needed to
evaluate the conclusions in the article are present in the article and/or the
Supplementary Materials. Additional data related to this paper may be requested
from the authors upon request.
SUPPLEMENTARY MATERIAL
SUMMARY
section S1. Piezoelectric transducer design using the KLM model
section S2. SAF imaging and DMAS algorithm
fig. S1. Testing performance of a commercial rigid probe on curved surfaces.
fig. S2. Schematic illustration of the device fabrication process.
fig. S3. Ecoflex thickness as a function of spin coating speed on a glass slide.
fig. S4. Acoustic damping effects of silicone substrates.
fig. S5. The vibration mode comparison between PZT and 1-3 composites.
fig. S6. Pulse-echo response and bandwidth differences of transducers with and
without the backing layer (KLM simulation).
fig. S7. Bottom electrode design.
fig. S8. Top electrode design.
fig. S9. Four-layer top electrode fabrication processes.
fig. S10. Optical images of Cu serpentine interconnections under different laser
parameters.
fig. S11. Laser ablation resolution experiments.
fig. S12. Photographs of the device seamlessly laminated on different curved
surfaces.
fig. S13. ACF cable bonding.
fig. S14. Simulation results from the KLM model.
fig. S15. Dielectric properties of the device.
fig. S16. The phase angle change during the fabrication process and after
repetitive testing.
fig. S17. Experimental and simulation of a small array under biaxial tensile
strain.
fig. S18. Electric impedances under different bending curvatures.
fig. S19. The real and imaginary parts of electrical impedance under different
levels of bending and stretching.
fig. S20. Relative resistance changes of Cu serpentine under stretching.
fig. S21. Instruments for nondestructive evaluation.
fig. S22. Switch circuit of the entire testing system.
fig. S23. Reconstructed images based on simulation under flat, concave, and
convex surfaces.
fig. S24. The pulse-echo signal and 2D image of the two defects.
fig. S25. Polarization conditions.
fig. S26. The matching circuit of the ultrasound testing system.
fig. S27. Ultrasound signal filtering.
fig. S28. Simplified schematics of a transducer element.
fig. S29. The electrical model of a transducer.
fig. S30. General diagram showing the transmission line model of a two-port
system.
fig. S31. Schematics showing the basic concept of SAF.
fig. S32. Block diagrams for the imaging algorithms.
table S1. Parameters for the 1-3 composite, backing layer, and Ecoflex.
movie S1. Simulation of wave field under a planar surface.
movie S2. Simulation of wave field under a concave surface.
movie S3. Simulation of wave field under a convex surface.
Reference (59)
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INFORMATION & AUTHORS
InformationAuthors
INFORMATION
PUBLISHED IN
Science Advances
Volume 4 | Issue 3
March 2018
COPYRIGHT
Copyright © 2018 The Authors, some rights reserved; exclusive licensee American
Association for the Advancement of Science. No claim to original U.S. Government
Works. Distributed under a Creative Commons Attribution NonCommercial License
4.0 (CC BY-NC).
This is an open-access article distributed under the terms of the Creative
Commons Attribution-NonCommercial license, which permits use, distribution, and
reproduction in any medium, so long as the resultant use is not for commercial
advantage and provided the original work is properly cited.
SUBMISSION HISTORY
Received: 3 November 2017
Accepted: 8 February 2018
PERMISSIONS
See the Reprints and Permissions page for information about permissions for this
article.
ACKNOWLEDGMENTS
We thank S. Cai’s group at the University of California San Diego (UCSD) for the
Young’s modulus testing, Y. Dai for analyzing the strain distribution of the
device in stretching states, and S. Xiang for constructive feedback on the
manuscript preparation. Funding: The project described was partially supported
by the NIH (grant R21EB025521) and Clinical and Translational Science Awards
funding (grant UL1TR001442). Additional partial support was provided by the UCSD
Center for Healthy Aging, the U.S. Federal Railroad Administration through grant
FR-RRD-0027-11, and the U.S. NSF through grant CMMI-1362144. The content is
solely the responsibility of the authors and does not necessarily represent the
official views of the NIH or NSF. Author contributions: H.H., X.Z., Chonghe
Wang, L.Z., X.L., F.L.d.S., and S.X. designed and executed the experiments,
reconstructed images, wrote the paper, and were responsible for Figs. 2 (A, B,
F, and G), 3, 4, and 5. S.L. and M.P. carried out FEA simulations and were
responsible for Fig. 2E. R.C., Z.C., and Q.Z. characterized the
electromechanical coupling performances and were responsible for Fig. 2 (C and
D). Z.H., Chunfeng Wang, and Y. Gu edited the manuscript, characterized
mechanical properties, and were responsible for Fig. 1 (F to H). Y.C., Y. Lei,
and T.Z. were responsible for Fig. 1 (C to E). Chonghe Wang and N.K. were
responsible for Fig. 1 (A and B). Y. Guo, W.Z., Y. Li, and A.N. made
contributions in device design and fabrication. Y.T. and S.S. made contributions
in algorithm development. Competing interests: S.X., L.Z., Chonghe Wang, H.H.,
and X.L. are inventors on a Patent Cooperation Treaty (PCT) patent application
related to this work filed by the Regents of the University of California
(PCT/IS18/13116, filed 10 January 2018). The other authors declare that they
have no competing interests. Data and materials availability: All data needed to
evaluate the conclusions in the article are present in the article and/or the
Supplementary Materials. Additional data related to this paper may be requested
from the authors upon request.
AUTHORS
AFFILIATIONSEXPAND ALL
HONGJIE HU* HTTPS://ORCID.ORG/0000-0002-3792-4894
Materials Science and Engineering Program, University of California San Diego,
La Jolla, CA 92093–0418, USA.
View all articles by this author
XUAN ZHU*
Department of Structural Engineering, University of California San Diego, La
Jolla, CA 92161, USA.
View all articles by this author
CHONGHE WANG* HTTPS://ORCID.ORG/0000-0002-1083-2666
Department of NanoEngineering, University of California San Diego, La Jolla, CA
92093–0448, USA.
View all articles by this author
LIN ZHANG* HTTPS://ORCID.ORG/0000-0002-0230-5911
Department of NanoEngineering, University of California San Diego, La Jolla, CA
92093–0448, USA.
View all articles by this author
XIAOSHI LI
Materials Science and Engineering Program, University of California San Diego,
La Jolla, CA 92093–0418, USA.
View all articles by this author
SEUNGHYUN LEE HTTPS://ORCID.ORG/0000-0002-1354-3261
Department of Mechanical Engineering, Texas A&M University, College Station, TX
77843, USA.
View all articles by this author
ZHENLONG HUANG HTTPS://ORCID.ORG/0000-0002-3410-1411
Department of NanoEngineering, University of California San Diego, La Jolla, CA
92093–0448, USA.
State Key Laboratory of Electronic Thin Films and Integrated Devices, University
of Electronic Science and Technology of China, Chengdu, Sichuan 610054, P. R.
China.
View all articles by this author
RUIMIN CHEN HTTPS://ORCID.ORG/0000-0002-5338-359X
Department of Ophthalmology and Department of Biomedical Engineering, University
of Southern California, Los Angeles, CA 90089, USA.
View all articles by this author
ZEYU CHEN
Department of Ophthalmology and Department of Biomedical Engineering, University
of Southern California, Los Angeles, CA 90089, USA.
View all articles by this author
CHUNFENG WANG HTTPS://ORCID.ORG/0000-0001-9676-8425
Department of NanoEngineering, University of California San Diego, La Jolla, CA
92093–0448, USA.
The Key Laboratory of Materials Processing and Mold of Ministry of Education,
School of Materials Science and Engineering, School of Physics and Engineering,
Zhengzhou University, Zhengzhou, Henan 450001, P. R. China.
View all articles by this author
YUE GU HTTPS://ORCID.ORG/0000-0003-0437-5339
Materials Science and Engineering Program, University of California San Diego,
La Jolla, CA 92093–0418, USA.
View all articles by this author
YIMU CHEN
Department of NanoEngineering, University of California San Diego, La Jolla, CA
92093–0448, USA.
View all articles by this author
YUSHENG LEI
Department of NanoEngineering, University of California San Diego, La Jolla, CA
92093–0448, USA.
View all articles by this author
TIANJIAO ZHANG HTTPS://ORCID.ORG/0000-0002-9758-4895
Materials Science and Engineering Program, University of California San Diego,
La Jolla, CA 92093–0418, USA.
View all articles by this author
NAMHEON KIM HTTPS://ORCID.ORG/0000-0002-3614-2233
Department of NanoEngineering, University of California San Diego, La Jolla, CA
92093–0448, USA.
View all articles by this author
YUXUAN GUO HTTPS://ORCID.ORG/0000-0001-7582-4555
Department of NanoEngineering, University of California San Diego, La Jolla, CA
92093–0448, USA.
View all articles by this author
YUE TENG HTTPS://ORCID.ORG/0000-0002-3032-3834
Department of Mathematics, University of California San Diego, La Jolla, CA
92093, USA.
View all articles by this author
WENBO ZHOU HTTPS://ORCID.ORG/0000-0001-5980-3287
Department of Physics, University of California San Diego, La Jolla, CA 92093,
USA.
View all articles by this author
YANG LI
Department of NanoEngineering, University of California San Diego, La Jolla, CA
92093–0448, USA.
View all articles by this author
AKIHIRO NOMOTO
Department of NanoEngineering, University of California San Diego, La Jolla, CA
92093–0448, USA.
View all articles by this author
SIMONE STERNINI HTTPS://ORCID.ORG/0000-0002-6511-2666
Department of Structural Engineering, University of California San Diego, La
Jolla, CA 92161, USA.
View all articles by this author
QIFA ZHOU
Department of Ophthalmology and Department of Biomedical Engineering, University
of Southern California, Los Angeles, CA 90089, USA.
View all articles by this author
MATT PHARR HTTPS://ORCID.ORG/0000-0001-8738-5393
Department of Mechanical Engineering, Texas A&M University, College Station, TX
77843, USA.
View all articles by this author
FRANCESCO LANZA DI SCALEA
Department of Structural Engineering, University of California San Diego, La
Jolla, CA 92161, USA.
View all articles by this author
SHENG XU† HTTPS://ORCID.ORG/0000-0002-3120-4992 SHENGXU@UCSD.EDU
Materials Science and Engineering Program, University of California San Diego,
La Jolla, CA 92093–0418, USA.
Department of NanoEngineering, University of California San Diego, La Jolla, CA
92093–0448, USA.
View all articles by this author
FUNDING INFORMATION
National Institutes of Health: UL1TR001442
National Institutes of Health: R21EB025521
US National Science Foundation: CMMI-1362144
US Federal Railroad Administration: FR-RRD-0027-11
UC San Diego Center for Healthy Aging
NOTES
*
These authors contributed equally to this work.
†
Corresponding author. Email: shengxu@ucsd.edu
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MEDIA
FiguresMultimedia
FIGURES
Fig. 1 Schematics and design of the stretchable ultrasonic transducer array.
(A) Schematics showing the device structure. (B) Exploded view to illustrate
each component in an element. (C) The optical image (bottom view) of four
elements, showing the morphology of the piezoelectric material and bottom
electrodes. (D) The tilted scanning electron microscopy image of a 1-3
piezoelectric composite. (E) The optical image (top view) of four elements,
showing the morphology of the backing layer and top electrodes. (F to H) Optical
images of this stretchable device when (F) bent around a developable surface,
(G) wrapped on a nondevelopable surface, and (H) in a mixed mode of folding,
stretching, and twisting, showing its mechanical robustness.
GO TO FIGUREOPEN IN VIEWER
Fig. 2 Characterizations of piezoelectric and mechanical properties.
(A) The impedance and phase angle spectra of the 1-3 composite before and after
processing, showing good electromechanical coupling of the fabricated transducer
(keff, ~0.60; θ, ~50°). (B) Pulse-echo response and frequency spectra, with a
short spatial pulse length (~1.94 μs), a high SNR (~20.24 dB), and a wide
bandwidth (~47.11%). (C) The resonance and antiresonance frequency variations of
the 100 transducer elements. The mean values/SDs are 3.51 MHz/56.8 kHz
(resonant) and 4.30 MHz/59.1 kHz (antiresonant), respectively. The 100% yield
demonstrates fabrication robustness. (D) Average cross-talk levels between
elements that are adjacent, two elements away, and three elements away, showing
the outstanding anti-interference capacity of the device. (E) The optical image
(left) and corresponding finite element analysis (FEA) simulation (right) of a 2
× 2 array under 50% biaxial tensile strain, showing its excellent
stretchability. The local strain level (maximum principal strain) in the
interconnects is indicated by the color scale. (F) The optical image after
releasing the biaxial strain of 50%. The zoomed-in image highlights plastic
deformation and local delamination of the interconnects upon loading/unloading.
(G) Electrical impedances of the transducer under different strain levels,
showing the mechanical stability of the device.
GO TO FIGUREOPEN IN VIEWER
Fig. 3 Characterization of spatial resolution.
(A) Schematics of spatial resolution measurement setup, with focal lengths of
20, 32, 37, and 52 mm, respectively. (B) Comparison of noise floors
reconstructed by DMAS and DAS algorithms, revealing the benefits of the DMAS
algorithm with only 0.01% energy ratio of noise to the reflector. (C) Images of
wire phantom combining the four tests with different f numbers, showing the
capability of focusing at different depths and obtaining high-resolution images.
(D) Axial and (E) lateral line spread functions for the center wire at different
focal lengths. Resolution is defined as the linespread function width at an
intensity of −6 dB. (F) Experimental (Exp.) and simulation (Simu.) results of
lateral and axial resolutions.
GO TO FIGUREOPEN IN VIEWER
Fig. 4 Two-dimensional images of a linear defect under complex surfaces.
Optical images of experimental setups with the stretchable ultrasonic device
tested on (A) planar, (B) concave, and (C) convex surfaces showing the good
conformability of the device on these surfaces (first column); simulation
results showing the different wave fields and sensing modes (second column);
pulse-echo signals from the defects and boundaries with high SNR (third column);
and acquired 2D images using DMAS algorithms with accurate and artifact-free
positions (fourth column). S wave, shear wave; L wave, longitudinal wave; SNRD,
SNR of pulse-echo response from the defect.
GO TO FIGUREOPEN IN VIEWER
Fig. 5 Three-dimensional image reconstruction of intricate defects under a
convex surface.
(A) Schematics of the experimental setup, illustrating the spatial location and
relative orientation of the two defects in the test subject. (B) The
reconstructed 3D image, showing complete geometries of the two defects. (C to E)
The 3D image from different view angles, showing the relative positions and
orientations of the two defects to the top surface, which match the design well.
GO TO FIGUREOPEN IN VIEWER
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FiguresTables
View figure
Fig. 1
Fig. 1 Schematics and design of the stretchable ultrasonic transducer array.
(A) Schematics showing the device structure. (B) Exploded view to illustrate
each component in an element. (C) The optical image (bottom view) of four
elements, showing the morphology of the piezoelectric material and bottom
electrodes. (D) The tilted scanning electron microscopy image of a 1-3
piezoelectric composite. (E) The optical image (top view) of four elements,
showing the morphology of the backing layer and top electrodes. (F to H) Optical
images of this stretchable device when (F) bent around a developable surface,
(G) wrapped on a nondevelopable surface, and (H) in a mixed mode of folding,
stretching, and twisting, showing its mechanical robustness.
View figure
Fig. 2
Fig. 2 Characterizations of piezoelectric and mechanical properties.
(A) The impedance and phase angle spectra of the 1-3 composite before and after
processing, showing good electromechanical coupling of the fabricated transducer
(keff, ~0.60; θ, ~50°). (B) Pulse-echo response and frequency spectra, with a
short spatial pulse length (~1.94 μs), a high SNR (~20.24 dB), and a wide
bandwidth (~47.11%). (C) The resonance and antiresonance frequency variations of
the 100 transducer elements. The mean values/SDs are 3.51 MHz/56.8 kHz
(resonant) and 4.30 MHz/59.1 kHz (antiresonant), respectively. The 100% yield
demonstrates fabrication robustness. (D) Average cross-talk levels between
elements that are adjacent, two elements away, and three elements away, showing
the outstanding anti-interference capacity of the device. (E) The optical image
(left) and corresponding finite element analysis (FEA) simulation (right) of a 2
× 2 array under 50% biaxial tensile strain, showing its excellent
stretchability. The local strain level (maximum principal strain) in the
interconnects is indicated by the color scale. (F) The optical image after
releasing the biaxial strain of 50%. The zoomed-in image highlights plastic
deformation and local delamination of the interconnects upon loading/unloading.
(G) Electrical impedances of the transducer under different strain levels,
showing the mechanical stability of the device.
View figure
Fig. 3
Fig. 3 Characterization of spatial resolution.
(A) Schematics of spatial resolution measurement setup, with focal lengths of
20, 32, 37, and 52 mm, respectively. (B) Comparison of noise floors
reconstructed by DMAS and DAS algorithms, revealing the benefits of the DMAS
algorithm with only 0.01% energy ratio of noise to the reflector. (C) Images of
wire phantom combining the four tests with different f numbers, showing the
capability of focusing at different depths and obtaining high-resolution images.
(D) Axial and (E) lateral line spread functions for the center wire at different
focal lengths. Resolution is defined as the linespread function width at an
intensity of −6 dB. (F) Experimental (Exp.) and simulation (Simu.) results of
lateral and axial resolutions.
View figure
Fig. 4
Fig. 4 Two-dimensional images of a linear defect under complex surfaces.
Optical images of experimental setups with the stretchable ultrasonic device
tested on (A) planar, (B) concave, and (C) convex surfaces showing the good
conformability of the device on these surfaces (first column); simulation
results showing the different wave fields and sensing modes (second column);
pulse-echo signals from the defects and boundaries with high SNR (third column);
and acquired 2D images using DMAS algorithms with accurate and artifact-free
positions (fourth column). S wave, shear wave; L wave, longitudinal wave; SNRD,
SNR of pulse-echo response from the defect.
View figure
Fig. 5
Fig. 5 Three-dimensional image reconstruction of intricate defects under a
convex surface.
(A) Schematics of the experimental setup, illustrating the spatial location and
relative orientation of the two defects in the test subject. (B) The
reconstructed 3D image, showing complete geometries of the two defects. (C to E)
The 3D image from different view angles, showing the relative positions and
orientations of the two defects to the top surface, which match the design well.
Reference #1