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Flexible SU-8 microstructures for neural implant design
Available online at www.sciencedirect.com
Sensors and Actuators A 147 (2008) 324–331
Flexible SU-8 microstructures for neural implant design夽 J.P.F. Spratley, M.C.L. Ward ∗ , P.S. Hall, C. Thursfield School of Engineering, The University of Birmingham, Birmingham, UK Received 8 June 2007; received in revised form 20 August 2007; accepted 12 October 2007 Available online 25 October 2007
Abstract This work investigates the use of SU-8 microstructures for applications within flexible neural implants. Six different microstructures are fabricated and tested in both tension and bending to determine their Young’s modulus and their failure stress when deformed about a small radius of curvature. A numerical model is presented that accurately predicts the performance of the structures under bending, and physical testing shows that stresses of up to 300 MPa are achievable. © 2007 Elsevier B.V. All rights reserved. Keywords: MEMS; SU-8; Materials testing
1. Introduction A key issue in determining the quality of a person’s life is the degree to which they are able to exert control over their own destiny. Much of this control is manifest in our ability to communicate and to physically interact with our environment. Neurological conditions such as motor neurone disease, severe spinal injury, brain stem stroke or cerebral palsy often significantly preclude communication and physical interaction. Linked with intact intellect and cognition, this can lead to high levels of frustration and low quality of life. In many cases patients with these conditions have the prospect of many years of future life, even for degenerative neurological conditions because of the increasing practice of artificial ventilation. Current clinical practice attempts to identify residual, often minimal and in some cases, degenerating physical ability and to harness this by using simple switch activation as an input to a system. This can be used to facilitate communication or directly operate equipment within the environment, such as a powered wheelchair or an attention calling system. Due to the severe physical limitations it is often the case that only one or two reliable switch sites can be found and this falls well short of an adequate input regime to any system that, for instance, can produce communicative output at a rate that is consistent with effective social interaction.
夽 This paper is part of the Special Issue Transducers 07/Eurosensors XXI, Sensors and Actuators A: Physical; Volume 145–146C. ∗ Corresponding author. Tel.: +44 121 414 4172. E-mail address: [email protected] (M.C.L. Ward).
0924-4247/$ – see front matter © 2007 Elsevier B.V. All rights reserved. doi:10.1016/j.sna.2007.10.052
The authors have previously presented a design for a microsensor that can be injected into the motor cortex of the brain, allowing motor-signals to be transmitted wirelessly to a external processor [1]. This will give the user control over a computer, and therefore increased ability to communicate or interact with their surroundings, for example sending an email, controlling a motorised wheelchair, or simply turning on the lights. The design is shown in Fig. 1. The major components are: 1. The neural connector spikes; this array of spikes penetrate into the surface of the motor cortex to read the electrical signal from the surrounding groups of neurons. 2. The inductor coils; these coils allow wireless communication, which is essential to remove the infection paths that occur by passing wires through the skin and other boundaries within the body. 3. The wings; the wings provide mounting for the inductor coils and other electronic components, and maximize the communications potential of the device. To prevent physical damage to the brain the wings lie upon the surface of the motor cortex, within the brains protective dura. 4. The flexible beams; the wings are attached to the main body of the device by a flexible structure. This allows the wings to be folded through 90◦ and therefore pass through a standard cannula (inside diameter 1.3 mm). Making the device injectable reduces the severity of surgery and the associated risks. This paper will address the major mechanical challenge of this device, the design of the flexible beams. SU-8 has been
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2. Testing 2.1. Test piece design and fabrication
Fig. 1. Design of neural implant.
chosen as the base material used in this design as it: has been shown to be biocompatible [2,3]; is an insulator and therefore good for mounting electrical components [4,5], and has a relatively low Young’s modulus and therefore suitable for a highly flexible design [6–8]. SU-8 is a well-established material used within micromechanical processing. Its unique properties as a thick film negative photoresist, capable of high aspect ratios when exposed in the ‘near UV’ spectrum has been well documented [6–11]. It can be used as a sacrificial material for selective etching [12] or as a selfsacrificial layer [13,14], a mold for casting of other materials or electroplating [10,12,15,16], and as a structural material integral to the fabricated device [15,17]. This functionality has allowed SU-8 to be successfully applied to a wide variety of applications from microgrippers [17–19], to piezoelectric motor components [15], and microfluidic devices [20–22]. The basic mechanical properties of SU-8 are also well documented, however they are largely affected by processing parameters. Lorenz et al. [7] measured the Young’s modulus of SU-8 in a tensile test as 4.02 GPa and noted that there was no plastic domain observed. Hopcroft et al. [6] created SU-8 cantilevers, and under bending, reported the Young’s modulus to be 2–3 GPa. Chang et al. [8] conducted tests using Speckle interferometry with electron microscopy (SIEM) to determine the mechanical properties and determined the Young’s modulus to be 1.54–3.09 GPa. The ultimate tensile strength (UTS) has been reported by McAleavey [11] at 120–130 MPa and Chang et al. [8] at 49–77 MPa. This study will add to the above body of work by characterising SU-8 under conditions of high bending, and also will compare the performance of structures with different configurations of thin beams in both bending and tension. The processing parameters will be reported to allow for repeatability and will include a final hard-bake to remove any surface cracks normally associated with SU-8, therefore increasing the failure stress of the material. As the processing parameters vary the Young’s modulus this property will also be measured and, using the value gained, a mathematical model will be presented to accurately predict the performance under bending.
The test pieces were designed as a ‘dog-bone’ shape with a 4.5 mm × 1 mm flexible structure between two 5 mm × 5 mm paddles, as shown in Fig. 2. This design allows for micro-scale structures to make up the flexible structure, whilst the large ends make handling easy. In device A the flexible structure is solid, whereas in devices B through to F the structure consists of patterns of 25 m thick beams; in B the beams are all parallel to the length of the device spaced by 25 m, whereas devices C and D have perpendicular cross-links at 750 and 125 m spacings, respectively. In devices E and F the beams are at 45◦ to the length of the structure creating a square diamond mesh with distances between the beams of 75 m for E and 225 m for F. Fig. 3 shows SEM images of the devices once they have been released from the silicon wafer. The devices were fabricated from the 2000 series of SU-8 using cyclopentanone as a solvent rather than the gamma butyrolactone (GBL) used with the original series. The specific type of resist used is SU-8 2050 from MicroChem Corporation, using the following process to create 50 m thick structures (Fig. 4). The SU-8 2050 was spun onto a standard 4” 1 0 0 silicon wafer using a two-step spin cycle: 10 s at 500 rpm to spread the SU-8 evenly and then 30 s at 3250 rpm to achieve the required thickness. A soft-bake was performed on a hotplate for 3 min at 65 ◦ C ramping up to 95 ◦ C for a further 7 min and the wafer was removed from the hot plate and allowed to air cool. The SU-8 was selectively exposed using a Canon PLA501 Mask Aligner through a darkfield mask at 160 mJ/cm2 and then baked at 65 ◦ C for 1 min and ramped to 95 ◦ C for 6 min. The combination of the exposure and a post-exposure bake enables the material to be polymerised at the points of exposure (being a negative resist). The SU-8 was developed by immersion in EC Solvent for 6 min removing all the unexposed material. The wafer was then cleaned using IPA and distilled water, followed by blowdrying in a stream of dry nitrogen gas. The wafer was finally hard-baked at 200 ◦ C for 5 min. The hardbake anneals the SU-8 and removing and surface micro-cracks that may cause the structures to be structurally weak and prone to failure at low stresses. The difference between SU-8 structures with and without hard baking can be seen in Fig. 5. To release the SU-8 structures from the wafer they are immersed in a bath of a 20% sodium hydroxide (NaOH) solution at room temperature for approximately 2 h.
Fig. 2. Design of the ‘Dog-bone’ test piece (all dimensions in mm).
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Fig. 3. SEM images of released test pieces. (a) Device A; (b) device B; (c) device C; (d) device D; (e) device E; (f) device F.
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2.2. Test protocol The devices will be tested using two different methods; one to analyze their performance under tension and the other under high deformation bending.
Fig. 4. Fabrication process.
2.2.1. Tension Tensile testing was undertaken using a Bose ElectroForce 3200 test instrument, allowing displacements to a resolution of 1 m and forces to a resolution of 0.01 N to be measured. The end handles of the devices were securely adhered to metal plates, which were clamped in to the jaws of the test cell. The metal plates prevented any mechanical damage occurring to the devices due to the serrated jaws of the clamps. The machine stretched the devices at a rate of 10 m/s and recorded the applied load. The force-displacement data was converted into a stress–strain curve by applying the physical dimensions of the test piece to the following equation: E=
σ P/χ = L/L
(1)
where E is the Young’s modulus of the material in N/m2 ; σ the stress (N/m2 ); the strain, which is dimensionless; P is the applied force (N); χ the cross-sectional area (m2 ); L the change in length (m) and L is the original length (m). The cross-sectional area for A and B are constant, and therefore simple, however devices C through to F have changing cross sections, due to cross members in C and D and the diamond structure in E and F. Therefore in these cases the cross sectional area is averaged along the length of the flexible section. 2.2.2. Bending The authors have previously demonstrated a novel test procedure for applying large deformations in bending to highly flexible test structures [23]. A schematic of this test setup is shown in Fig. 6. The balance and clamp were placed below the load cell and the load cell moved downwards until it came into contact with the top of the clamp, which was taken as the datum position. The load cell was moved to a position 3 mm above the datum, and the test structure was placed in the clamp in a vertical orientation and the scales zeroed. The clamp and test structure were then slid underneath the load cell, imparting an initial curvature of 3 mm on the structure. The load cell was moved downwards at a speed of 1 mm/min until the test structure fractured or the
Fig. 5. SEM images showing the effect of hardbaking on SU-8 surface properties. (a) Before hardbake and (b) after hardbake.
Fig. 6. Schematic of test rig.
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load cell moved the full 3 mm. The position of the load cell and the reading from the scales were outputted to a PC to create a force–displacement reading for the structure. For this test method friction between the test piece and load cell needed to be minimised, and therefore the base of the load cell was ground to a very smooth finish. A force–displacement recording was also taken of the balance to remove any errors from the balance compressing as the force was applied, and this was subtracted from the readings to give an accurate result.
where φ0 is the angle of the free end. The equation was solved using a Runge–Kutta solver with φ0 set to π/2 and the force iterated until φ at ξ = 1 equaled zero. A loop in the programming was introduced to reduce the length of the beam in steps, thereby simulating increasingly tighter curves. The curve shapes are shown in Fig. 8. In this model the friction between the test piece and the load cell is modelled such that it creates a compressive force within the beam that is transmitted to the base as vertical force. Therefore the force transmitted to the balance becomes P :
3. Bending theory, numerical analysis
P = P + FR = P(1 + μ)
A numerical model has been created to predict the bending characteristics of the structures. Fig. 7 shows the orientation of the beam and force in this model. The equation for a cantilever beam is [24]:
where μ is the dimensionless coefficient of friction of SU-8, quoted by Lorenz et al. [7] as 0.19. The force–displacement curves can therefore be generated and directly compared with the physical results and those of the simple mathematical model. Given the standard engineering equation:
dφ 1 M(x) = (2) = EI R ds where φ is the angle of the beam in radians at arc length, s (m); M(x) the bending moment (Nm); I the second moment of area (m4 ); and R is the radius of curvature (m). The normalized form of the governing equation for a linear elastic beam with these support and loading conditions is[25]: PL2 d2 φ sin φ = − EI dξ 2
(3)
where ξ = s/L, and the boundary conditions are: φ = φ0 φ=0
and
dφ =0 dξ
at ξ = 0
at ξ = 1
Fig. 7. Deflection of vertical beam under vertical point force.
E σ = R y
(5)
(6)
the maximum stress, σmax , that will occur in the outer edges of the beam where the radius of curvature is at a minimum, Rmin . This can therefore can be determined by setting y = d/2, and using the numerical model to determine Rmin and therefore: σmax =
Ed 2Rmin
(7)
4. Results and discussion (4a) (4b)
The stress–strain plots for the tensile tests are shown in Fig. 9, with a trendline showing the Young’s modulus for the devices. As shown in Eq. 1 the cross-sectional area and initial lengths are required to calculate the stress and strain. For device A the area is simply calculated by b × d, whereas for device B the area is calculated as the sum of cross-sections of all the 25 m thick beams. For both A and B the original length is simply the length of the flexible structure, 4.54 mm. Devices C and D are not so simple to process; for both of these devices the cross-members
Fig. 8. Curve shapes of the beam radius created by the numerical model.
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Fig. 9. Tensile: stress–strain results.
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shorten the effective length of the flexible section, and therefore the length used is equal to the actual length minus the sum of all the 25 m cross-member widths. This equals 4.35 mm for C and 3.6 mm for D. The cross section used is the same as B (the sum of the longitudinal beam’s areas). As can be seen from the results devices A–D all confirm a Young’s modulus of 2 GPa. What is noticeable from the graph is the methods and points of failure of these structures; devices A and D perform in a linear manner and fail catastrophically with A under the highest stress of 77 MPa. In contrast B and C undergo a gradual degradation as their beams do not all break at the same time. Device C undergoes the greatest strain of 4.6%, and device F fails at the least strain, 2.4%. Devices E and F exhibit significantly different results from A to D due to their geometry. The mechanism of failure will be in three parts: firstly as the initial load is applied the outside longitudinal beams will stretch and fail; next the diamond structure
Fig. 10. Bending: force against bend radius results. (a) Device A; (b) device B; (c) device C; (d) device D; (e) device E; (f) device F.
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will start to collapse lengthening the beam without significantly stressing the material, and finally the complete structure of the diamond lattice will fail due to a combination of the stress imparted by the twisting of the collapsing structure and the tension applied by the longitudinal force. This ‘concertina’ failure reduced the appearance of the Young’s modulus of the structure to 0.5 GPa. Fig. 10 shows the results of the bending test for three of each of the test devices. Each graph also contains the curve calculated from the numerical model. The values used in the equations are: E = 2 GPa (from the tensile test results); d = 43 m measured from the fabricated test pieces and μ = 0.19. For devices A and B, the second moment of area is calculated by the standard equation: I=
bd 3 12
(8)
for devices C and D, I is calculated by averaging the second moment of area of the longitudinal beams against that for the cross members, with the ratio according to the total of their respective lengths. For devices E and F the second moment of area is constantly changing along the length of the beam, and therefore for simplicity the average value is used. It can be seen from the graphs that the model for devices A to D very accurately match the results gained from the physical testing. The models for E and F, although still being close to the tested results, are not quite as accurate. This is likely to be a result of the model simplifying the second moment of area for these complex structures. Fig. 11 shows the maximum stress in each of the beams at the point of failure, calculated by Eq. 7. Devices A and B perform very similarly, with one of the devices reaching a high maximum value of approximately 300 MPa and one of the group failing at a low stress of 50–150 MPa. For device A this is likely to be due to the presence of a surface defect causing crack initiation; this device is more susceptible to defects as it has a larger surface area at the point of maximum stress (where y = d/2). The failure of device B is less likely to have been due to a surface defect as its beams are unconnected and therefore one of them failing prematurely would not have caused the others to fail. It is more likely to be due to the instability of its long thin beams caused an increase in the stress due to uncontrolled twisting or splaying.
Fig. 11. Bending: maximum stress at failure.
Device C was much more consistent than either A or B, as it has neither the high surface area of A or the instability of B. Devices D–F all performed similarly and failed consistently at between 100 and 150 MPa. Devices D and E both have small square holes that would probably acted as stress concentrators in the corners. Likewise device F, albeit having much larger holes, has sharp corners that would have focused the stress and caused premature failure. 5. Conclusion This study has shown that it is possible to maximise the flexibility of MEMS planar structures with clever design. A numerical model has been put forward that accurately predicts the performance of the structures with constant beam shapes, or those that can be approximated as constant. Out of the devices tested the results suggest that the best design for the application is Device C as it minimises the problems that the other designs are affected by: it is less susceptible to defects as it has a small surface area at the point of maximum stress; it’s longitudinal beams have cross-members to decrease instability, however are not too close to act as stress concentrators, and when it fails it does so in a non-catastrophic manner. With this design it has been shown that it is possible to repeatably achieve a failure stress of 200–300 MPa in bending, which is much greater than the UTS that has been reported for tensile testing of 49–77 MPa [8] and 120–130 MPa [11]. References [1] J.P.F. Spratley, M.C.L. Ward, P.S. Hall, P.S.C. Thursfield, Design of an injectible microsensor to wirelessly transmit signals from the motor cortex of the human brain, in: The IET Seminar on MEMS Sensors and Actuators, Savoy Place, London, UK, April 28, 2006. [2] G. Voskerician, M.S. Shive, R.S. Shawgo, H. von Recum, J.M. Anderson, M.J. Cima, R. Langer, Biocompatibility and biofouling of mems drug delivery devices, Biomaterials 24 (11) (2003) 1959–1967. [3] G. Kotzara, M. Freasa, P. Abelb, A. Fleischmanc, S. Royc, C. Zormand, J.M. Morane, J. Melzak, Evaluation of mems materials of construction for implantable medical devices, Biomaterials 23 (3/4) (2002) 2737–2750. [4] R. Ramachandran, A.V.H. Pham, Development of rf/microwave on-chip inductors using an organic micromachining process, IEEE Trans. Adv. Pack. 25 (2) (2002) 244–247. [5] T. Kohlmeier, H.H. Gatzen, Challenges in using photosensitive embedding material to planarize multi-layer coils for actuator systems, J. Magn. Magn. Mater. 242–245 (2) (2002) 1149–1152. [6] M. Hopcroft, T. Kramer, G. Kim, K. Takashima, Y. Higo, D. Moore, J. Brugger, Micromechanical testing of su-8 cantilevers, Fatigue Fract. Eng. Mater. Struct. 28 (8) (2005) 735–742. [7] H. Lorenz, M. Despont, N. Fahrni, N. LaBianca, P. Renaud, P. Vettiger, Su-8: a low-cost negative resist for mems, J. Micromech. Microeng. 7 (3) (1997) 121–124. [8] S. Chang, J. Warren, F. Chiang, Mechanical testing of epon su-8 with seim, in: Microscale Systems: Mechanics and Measurements Syposium, Orlando, FL, US, 2000. [9] H. Lorenz, M. Laudon, P. Renaud, Mechanical characterization of a new high-aspect-ratio near UV-photoresist, Microelectron. Eng. 42 (1998) 371–374. [10] H. Lorenz, M. Despont, N. Fahrni, J. Brugger, P. Vettiger, P. Renaud, High-aspect-ratio, ultrathick, negative-tone near-UV photoresist and its applications for mems, Sens. Actuators A: Phys. 64 (1) (1998) 33–39.
J.P.F. Spratley et al. / Sensors and Actuators A 147 (2008) 324–331 [11] A. McAleavey, Mechanical properties of su-8. Materials science of microelectromechanical systems (MEMS) devices, in: Proceedings of the Symposium, Boston, MA, United States, December 1–2, 1998. [12] K.Y. Lee, N. LaBianca, S.A. Rishton, S. Zolgharnain, J.D. Gelorme, J. Shaw, T.H.P. Chang, Micromachining applications of a high resolution ultrathick photoresis, J. Vac. Sci. Technol. B 13 (6) (1995) 3012– 3016. [13] I.G. Foulds, M. Parameswaran, A planar self-sacrificial multilayer su-8based mems process utilizing a UV-blocking layer for the creation of freely moving parts, J. Micromech. Microeng. 16 (10) (2006) 2109–2115. [14] Z. Ling, K. Lian, In situ fabrication of su-8 movable parts by using pagdiluted su-8 as the sacrificial layer, Microsyst. Technol. Nanosyst. Stor. Process. Syst. 13 (3/4) (2007) 253–257. [15] L. Dellmann, S. Roth, C. Beuret, G.A. Racine, H. Lorenz, M. Despont, P. Renaud, P. Vettiger, N.F. deRooij, Fabrication process of high aspect ratio elastic structures for piezoelectric motor applications, in: Transducers 97—1997 International Conference on Solid-State Sensors and Actuators, 1997, pp. 641–644. [16] J.H. Li, D. Chen, J.Y. Zhang, J.Q. Liu, J. Zhu, Indirect removal of su-8 photoresist using pdms technique, Sens. Actuators A: Phys. 125 (2) (2006) 586–589. [17] H.S. Choi, D.C. Lee, S.S. Kim, C.S. Han, The development of a microgripper with a perturbation-based configuration design method, J. Micromech. Microeng. 15 (6) (2005) 1327–1333. [18] N.T. Nguyen, S.S. Ho, C.L.N. Low, A polymeric microgripper with integrated thermal actuators, J. Micromech. Microeng. 7 (14) (2004) 969–974. [19] I. Roch, P. Bidaud, D. Collard, L. Buchaillot, Fabrication and characterization of an su-8 gripper actuated by a shape memory alloy thin film, J. Micromech. Microeng. 13 (2) (2003) 330–336. [20] P. Abgrall, C. Lattes, V. Conederal, X. Dollat, S. Colin, A.M. Gue, A novel fabrication method of flexible and monolithic 3D microfluidic structures using lamination of su-8 films, J. Micromech. Microeng. 16 (2006) 113–121. [21] B.L. Gray, S. Jaffer, Mechanical and microfluidic interconnects fabricated in su-8 photopolymer, pp. 585–591, in: ICMM 2005: 3rd International Conference on Microchannels and Minichannels, Pt B, 2005. [22] H. Sato, H. Matsumura, S. Keino, S. Shoji, An all su-8 microfluidic chip with built-in 3D fine microstructures, J. Micromech. Microeng. 16 (11) (2006) 2318–2322. [23] J.P.F. Spratley, M.C.L. Ward, P.S. Hall, Bending characteristics of su-8, Micro Nano Lett. 2 (2) (2007) 20–23.
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Biographies Jon Spratley obtained his MEng in Mechanical Engineering in 2004, and is currently in his third year of studying for his PhD in the field of Microsystems and Nanotechnology at the University of Birmingham. His research interests are in neural and brain interfacing, and he is currently researching microstructure design for implantable systems. Dr. Ward obtained his physics degree from Imperial College in 1981 and subsequently was awarded his PhD from Warwick University for his EXAFS study on the passivation of stainless steels. In 1985 he joined RSRE Malvern where he worked for IR detectors, image processing, SOI electronics and MEMS before joining the University of Birmingham in 1999. He is currently senior lecture in microsystems engineering with research interest in noise, non-linear mechanics and microfluidics. Peter Hall is Professor of Communications Engineering, leader of the Antennas and Applied Electromagnetics Laboratory, and Head of the Devices and Systems Research Centre in the Department of Electronic, Electrical and Computer Engineering at The University of Birmingham. After graduating with a PhD in antenna measurements from Sheffield University, he spent 3 years with Marconi Space and Defence Systems, Stanmore working largely on a European Communications satellite project. He then joined The Royal Military College of Science as a Senior Research Scientist, progressing to Reader in Electromagnetics. He joined The University of Birmingham in 1994. He has researched extensively in the areas of microwave antennas and associated components and antenna measurements. Clive Thursfield started his career with an honours degree in electronics followed by a PhD in Biomedical Engineering from Strathclyde University where his research was associated with non-invasive cardiac monitoring. He is currently the Head of Research for the West Midland Rehabilitation Centre (WMRC) and for South Birmingham PCT. He is also the head of a newly created Department of Health Care Science (within WMRC) whose remit will be to ensure the use of scientific approaches and innovative and emerging technologies across the whole spectrum of rehabilitation. Clive is a Chartered Engineer and a Chartered Scientist.
Sensors and Actuators A 147 (2008) 324–331
Flexible SU-8 microstructures for neural implant design夽 J.P.F. Spratley, M.C.L. Ward ∗ , P.S. Hall, C. Thursfield School of Engineering, The University of Birmingham, Birmingham, UK Received 8 June 2007; received in revised form 20 August 2007; accepted 12 October 2007 Available online 25 October 2007
Abstract This work investigates the use of SU-8 microstructures for applications within flexible neural implants. Six different microstructures are fabricated and tested in both tension and bending to determine their Young’s modulus and their failure stress when deformed about a small radius of curvature. A numerical model is presented that accurately predicts the performance of the structures under bending, and physical testing shows that stresses of up to 300 MPa are achievable. © 2007 Elsevier B.V. All rights reserved. Keywords: MEMS; SU-8; Materials testing
1. Introduction A key issue in determining the quality of a person’s life is the degree to which they are able to exert control over their own destiny. Much of this control is manifest in our ability to communicate and to physically interact with our environment. Neurological conditions such as motor neurone disease, severe spinal injury, brain stem stroke or cerebral palsy often significantly preclude communication and physical interaction. Linked with intact intellect and cognition, this can lead to high levels of frustration and low quality of life. In many cases patients with these conditions have the prospect of many years of future life, even for degenerative neurological conditions because of the increasing practice of artificial ventilation. Current clinical practice attempts to identify residual, often minimal and in some cases, degenerating physical ability and to harness this by using simple switch activation as an input to a system. This can be used to facilitate communication or directly operate equipment within the environment, such as a powered wheelchair or an attention calling system. Due to the severe physical limitations it is often the case that only one or two reliable switch sites can be found and this falls well short of an adequate input regime to any system that, for instance, can produce communicative output at a rate that is consistent with effective social interaction.
夽 This paper is part of the Special Issue Transducers 07/Eurosensors XXI, Sensors and Actuators A: Physical; Volume 145–146C. ∗ Corresponding author. Tel.: +44 121 414 4172. E-mail address: [email protected] (M.C.L. Ward).
0924-4247/$ – see front matter © 2007 Elsevier B.V. All rights reserved. doi:10.1016/j.sna.2007.10.052
The authors have previously presented a design for a microsensor that can be injected into the motor cortex of the brain, allowing motor-signals to be transmitted wirelessly to a external processor [1]. This will give the user control over a computer, and therefore increased ability to communicate or interact with their surroundings, for example sending an email, controlling a motorised wheelchair, or simply turning on the lights. The design is shown in Fig. 1. The major components are: 1. The neural connector spikes; this array of spikes penetrate into the surface of the motor cortex to read the electrical signal from the surrounding groups of neurons. 2. The inductor coils; these coils allow wireless communication, which is essential to remove the infection paths that occur by passing wires through the skin and other boundaries within the body. 3. The wings; the wings provide mounting for the inductor coils and other electronic components, and maximize the communications potential of the device. To prevent physical damage to the brain the wings lie upon the surface of the motor cortex, within the brains protective dura. 4. The flexible beams; the wings are attached to the main body of the device by a flexible structure. This allows the wings to be folded through 90◦ and therefore pass through a standard cannula (inside diameter 1.3 mm). Making the device injectable reduces the severity of surgery and the associated risks. This paper will address the major mechanical challenge of this device, the design of the flexible beams. SU-8 has been
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2. Testing 2.1. Test piece design and fabrication
Fig. 1. Design of neural implant.
chosen as the base material used in this design as it: has been shown to be biocompatible [2,3]; is an insulator and therefore good for mounting electrical components [4,5], and has a relatively low Young’s modulus and therefore suitable for a highly flexible design [6–8]. SU-8 is a well-established material used within micromechanical processing. Its unique properties as a thick film negative photoresist, capable of high aspect ratios when exposed in the ‘near UV’ spectrum has been well documented [6–11]. It can be used as a sacrificial material for selective etching [12] or as a selfsacrificial layer [13,14], a mold for casting of other materials or electroplating [10,12,15,16], and as a structural material integral to the fabricated device [15,17]. This functionality has allowed SU-8 to be successfully applied to a wide variety of applications from microgrippers [17–19], to piezoelectric motor components [15], and microfluidic devices [20–22]. The basic mechanical properties of SU-8 are also well documented, however they are largely affected by processing parameters. Lorenz et al. [7] measured the Young’s modulus of SU-8 in a tensile test as 4.02 GPa and noted that there was no plastic domain observed. Hopcroft et al. [6] created SU-8 cantilevers, and under bending, reported the Young’s modulus to be 2–3 GPa. Chang et al. [8] conducted tests using Speckle interferometry with electron microscopy (SIEM) to determine the mechanical properties and determined the Young’s modulus to be 1.54–3.09 GPa. The ultimate tensile strength (UTS) has been reported by McAleavey [11] at 120–130 MPa and Chang et al. [8] at 49–77 MPa. This study will add to the above body of work by characterising SU-8 under conditions of high bending, and also will compare the performance of structures with different configurations of thin beams in both bending and tension. The processing parameters will be reported to allow for repeatability and will include a final hard-bake to remove any surface cracks normally associated with SU-8, therefore increasing the failure stress of the material. As the processing parameters vary the Young’s modulus this property will also be measured and, using the value gained, a mathematical model will be presented to accurately predict the performance under bending.
The test pieces were designed as a ‘dog-bone’ shape with a 4.5 mm × 1 mm flexible structure between two 5 mm × 5 mm paddles, as shown in Fig. 2. This design allows for micro-scale structures to make up the flexible structure, whilst the large ends make handling easy. In device A the flexible structure is solid, whereas in devices B through to F the structure consists of patterns of 25 m thick beams; in B the beams are all parallel to the length of the device spaced by 25 m, whereas devices C and D have perpendicular cross-links at 750 and 125 m spacings, respectively. In devices E and F the beams are at 45◦ to the length of the structure creating a square diamond mesh with distances between the beams of 75 m for E and 225 m for F. Fig. 3 shows SEM images of the devices once they have been released from the silicon wafer. The devices were fabricated from the 2000 series of SU-8 using cyclopentanone as a solvent rather than the gamma butyrolactone (GBL) used with the original series. The specific type of resist used is SU-8 2050 from MicroChem Corporation, using the following process to create 50 m thick structures (Fig. 4). The SU-8 2050 was spun onto a standard 4” 1 0 0 silicon wafer using a two-step spin cycle: 10 s at 500 rpm to spread the SU-8 evenly and then 30 s at 3250 rpm to achieve the required thickness. A soft-bake was performed on a hotplate for 3 min at 65 ◦ C ramping up to 95 ◦ C for a further 7 min and the wafer was removed from the hot plate and allowed to air cool. The SU-8 was selectively exposed using a Canon PLA501 Mask Aligner through a darkfield mask at 160 mJ/cm2 and then baked at 65 ◦ C for 1 min and ramped to 95 ◦ C for 6 min. The combination of the exposure and a post-exposure bake enables the material to be polymerised at the points of exposure (being a negative resist). The SU-8 was developed by immersion in EC Solvent for 6 min removing all the unexposed material. The wafer was then cleaned using IPA and distilled water, followed by blowdrying in a stream of dry nitrogen gas. The wafer was finally hard-baked at 200 ◦ C for 5 min. The hardbake anneals the SU-8 and removing and surface micro-cracks that may cause the structures to be structurally weak and prone to failure at low stresses. The difference between SU-8 structures with and without hard baking can be seen in Fig. 5. To release the SU-8 structures from the wafer they are immersed in a bath of a 20% sodium hydroxide (NaOH) solution at room temperature for approximately 2 h.
Fig. 2. Design of the ‘Dog-bone’ test piece (all dimensions in mm).
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Fig. 3. SEM images of released test pieces. (a) Device A; (b) device B; (c) device C; (d) device D; (e) device E; (f) device F.
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2.2. Test protocol The devices will be tested using two different methods; one to analyze their performance under tension and the other under high deformation bending.
Fig. 4. Fabrication process.
2.2.1. Tension Tensile testing was undertaken using a Bose ElectroForce 3200 test instrument, allowing displacements to a resolution of 1 m and forces to a resolution of 0.01 N to be measured. The end handles of the devices were securely adhered to metal plates, which were clamped in to the jaws of the test cell. The metal plates prevented any mechanical damage occurring to the devices due to the serrated jaws of the clamps. The machine stretched the devices at a rate of 10 m/s and recorded the applied load. The force-displacement data was converted into a stress–strain curve by applying the physical dimensions of the test piece to the following equation: E=
σ P/χ = L/L
(1)
where E is the Young’s modulus of the material in N/m2 ; σ the stress (N/m2 ); the strain, which is dimensionless; P is the applied force (N); χ the cross-sectional area (m2 ); L the change in length (m) and L is the original length (m). The cross-sectional area for A and B are constant, and therefore simple, however devices C through to F have changing cross sections, due to cross members in C and D and the diamond structure in E and F. Therefore in these cases the cross sectional area is averaged along the length of the flexible section. 2.2.2. Bending The authors have previously demonstrated a novel test procedure for applying large deformations in bending to highly flexible test structures [23]. A schematic of this test setup is shown in Fig. 6. The balance and clamp were placed below the load cell and the load cell moved downwards until it came into contact with the top of the clamp, which was taken as the datum position. The load cell was moved to a position 3 mm above the datum, and the test structure was placed in the clamp in a vertical orientation and the scales zeroed. The clamp and test structure were then slid underneath the load cell, imparting an initial curvature of 3 mm on the structure. The load cell was moved downwards at a speed of 1 mm/min until the test structure fractured or the
Fig. 5. SEM images showing the effect of hardbaking on SU-8 surface properties. (a) Before hardbake and (b) after hardbake.
Fig. 6. Schematic of test rig.
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load cell moved the full 3 mm. The position of the load cell and the reading from the scales were outputted to a PC to create a force–displacement reading for the structure. For this test method friction between the test piece and load cell needed to be minimised, and therefore the base of the load cell was ground to a very smooth finish. A force–displacement recording was also taken of the balance to remove any errors from the balance compressing as the force was applied, and this was subtracted from the readings to give an accurate result.
where φ0 is the angle of the free end. The equation was solved using a Runge–Kutta solver with φ0 set to π/2 and the force iterated until φ at ξ = 1 equaled zero. A loop in the programming was introduced to reduce the length of the beam in steps, thereby simulating increasingly tighter curves. The curve shapes are shown in Fig. 8. In this model the friction between the test piece and the load cell is modelled such that it creates a compressive force within the beam that is transmitted to the base as vertical force. Therefore the force transmitted to the balance becomes P :
3. Bending theory, numerical analysis
P = P + FR = P(1 + μ)
A numerical model has been created to predict the bending characteristics of the structures. Fig. 7 shows the orientation of the beam and force in this model. The equation for a cantilever beam is [24]:
where μ is the dimensionless coefficient of friction of SU-8, quoted by Lorenz et al. [7] as 0.19. The force–displacement curves can therefore be generated and directly compared with the physical results and those of the simple mathematical model. Given the standard engineering equation:
dφ 1 M(x) = (2) = EI R ds where φ is the angle of the beam in radians at arc length, s (m); M(x) the bending moment (Nm); I the second moment of area (m4 ); and R is the radius of curvature (m). The normalized form of the governing equation for a linear elastic beam with these support and loading conditions is[25]: PL2 d2 φ sin φ = − EI dξ 2
(3)
where ξ = s/L, and the boundary conditions are: φ = φ0 φ=0
and
dφ =0 dξ
at ξ = 0
at ξ = 1
Fig. 7. Deflection of vertical beam under vertical point force.
E σ = R y
(5)
(6)
the maximum stress, σmax , that will occur in the outer edges of the beam where the radius of curvature is at a minimum, Rmin . This can therefore can be determined by setting y = d/2, and using the numerical model to determine Rmin and therefore: σmax =
Ed 2Rmin
(7)
4. Results and discussion (4a) (4b)
The stress–strain plots for the tensile tests are shown in Fig. 9, with a trendline showing the Young’s modulus for the devices. As shown in Eq. 1 the cross-sectional area and initial lengths are required to calculate the stress and strain. For device A the area is simply calculated by b × d, whereas for device B the area is calculated as the sum of cross-sections of all the 25 m thick beams. For both A and B the original length is simply the length of the flexible structure, 4.54 mm. Devices C and D are not so simple to process; for both of these devices the cross-members
Fig. 8. Curve shapes of the beam radius created by the numerical model.
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Fig. 9. Tensile: stress–strain results.
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shorten the effective length of the flexible section, and therefore the length used is equal to the actual length minus the sum of all the 25 m cross-member widths. This equals 4.35 mm for C and 3.6 mm for D. The cross section used is the same as B (the sum of the longitudinal beam’s areas). As can be seen from the results devices A–D all confirm a Young’s modulus of 2 GPa. What is noticeable from the graph is the methods and points of failure of these structures; devices A and D perform in a linear manner and fail catastrophically with A under the highest stress of 77 MPa. In contrast B and C undergo a gradual degradation as their beams do not all break at the same time. Device C undergoes the greatest strain of 4.6%, and device F fails at the least strain, 2.4%. Devices E and F exhibit significantly different results from A to D due to their geometry. The mechanism of failure will be in three parts: firstly as the initial load is applied the outside longitudinal beams will stretch and fail; next the diamond structure
Fig. 10. Bending: force against bend radius results. (a) Device A; (b) device B; (c) device C; (d) device D; (e) device E; (f) device F.
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will start to collapse lengthening the beam without significantly stressing the material, and finally the complete structure of the diamond lattice will fail due to a combination of the stress imparted by the twisting of the collapsing structure and the tension applied by the longitudinal force. This ‘concertina’ failure reduced the appearance of the Young’s modulus of the structure to 0.5 GPa. Fig. 10 shows the results of the bending test for three of each of the test devices. Each graph also contains the curve calculated from the numerical model. The values used in the equations are: E = 2 GPa (from the tensile test results); d = 43 m measured from the fabricated test pieces and μ = 0.19. For devices A and B, the second moment of area is calculated by the standard equation: I=
bd 3 12
(8)
for devices C and D, I is calculated by averaging the second moment of area of the longitudinal beams against that for the cross members, with the ratio according to the total of their respective lengths. For devices E and F the second moment of area is constantly changing along the length of the beam, and therefore for simplicity the average value is used. It can be seen from the graphs that the model for devices A to D very accurately match the results gained from the physical testing. The models for E and F, although still being close to the tested results, are not quite as accurate. This is likely to be a result of the model simplifying the second moment of area for these complex structures. Fig. 11 shows the maximum stress in each of the beams at the point of failure, calculated by Eq. 7. Devices A and B perform very similarly, with one of the devices reaching a high maximum value of approximately 300 MPa and one of the group failing at a low stress of 50–150 MPa. For device A this is likely to be due to the presence of a surface defect causing crack initiation; this device is more susceptible to defects as it has a larger surface area at the point of maximum stress (where y = d/2). The failure of device B is less likely to have been due to a surface defect as its beams are unconnected and therefore one of them failing prematurely would not have caused the others to fail. It is more likely to be due to the instability of its long thin beams caused an increase in the stress due to uncontrolled twisting or splaying.
Fig. 11. Bending: maximum stress at failure.
Device C was much more consistent than either A or B, as it has neither the high surface area of A or the instability of B. Devices D–F all performed similarly and failed consistently at between 100 and 150 MPa. Devices D and E both have small square holes that would probably acted as stress concentrators in the corners. Likewise device F, albeit having much larger holes, has sharp corners that would have focused the stress and caused premature failure. 5. Conclusion This study has shown that it is possible to maximise the flexibility of MEMS planar structures with clever design. A numerical model has been put forward that accurately predicts the performance of the structures with constant beam shapes, or those that can be approximated as constant. Out of the devices tested the results suggest that the best design for the application is Device C as it minimises the problems that the other designs are affected by: it is less susceptible to defects as it has a small surface area at the point of maximum stress; it’s longitudinal beams have cross-members to decrease instability, however are not too close to act as stress concentrators, and when it fails it does so in a non-catastrophic manner. With this design it has been shown that it is possible to repeatably achieve a failure stress of 200–300 MPa in bending, which is much greater than the UTS that has been reported for tensile testing of 49–77 MPa [8] and 120–130 MPa [11]. References [1] J.P.F. Spratley, M.C.L. Ward, P.S. Hall, P.S.C. Thursfield, Design of an injectible microsensor to wirelessly transmit signals from the motor cortex of the human brain, in: The IET Seminar on MEMS Sensors and Actuators, Savoy Place, London, UK, April 28, 2006. [2] G. Voskerician, M.S. Shive, R.S. Shawgo, H. von Recum, J.M. Anderson, M.J. Cima, R. Langer, Biocompatibility and biofouling of mems drug delivery devices, Biomaterials 24 (11) (2003) 1959–1967. [3] G. Kotzara, M. Freasa, P. Abelb, A. Fleischmanc, S. Royc, C. Zormand, J.M. Morane, J. Melzak, Evaluation of mems materials of construction for implantable medical devices, Biomaterials 23 (3/4) (2002) 2737–2750. [4] R. Ramachandran, A.V.H. Pham, Development of rf/microwave on-chip inductors using an organic micromachining process, IEEE Trans. Adv. Pack. 25 (2) (2002) 244–247. [5] T. Kohlmeier, H.H. Gatzen, Challenges in using photosensitive embedding material to planarize multi-layer coils for actuator systems, J. Magn. Magn. Mater. 242–245 (2) (2002) 1149–1152. [6] M. Hopcroft, T. Kramer, G. Kim, K. Takashima, Y. Higo, D. Moore, J. Brugger, Micromechanical testing of su-8 cantilevers, Fatigue Fract. Eng. Mater. Struct. 28 (8) (2005) 735–742. [7] H. Lorenz, M. Despont, N. Fahrni, N. LaBianca, P. Renaud, P. Vettiger, Su-8: a low-cost negative resist for mems, J. Micromech. Microeng. 7 (3) (1997) 121–124. [8] S. Chang, J. Warren, F. Chiang, Mechanical testing of epon su-8 with seim, in: Microscale Systems: Mechanics and Measurements Syposium, Orlando, FL, US, 2000. [9] H. Lorenz, M. Laudon, P. Renaud, Mechanical characterization of a new high-aspect-ratio near UV-photoresist, Microelectron. Eng. 42 (1998) 371–374. [10] H. Lorenz, M. Despont, N. Fahrni, J. Brugger, P. Vettiger, P. Renaud, High-aspect-ratio, ultrathick, negative-tone near-UV photoresist and its applications for mems, Sens. Actuators A: Phys. 64 (1) (1998) 33–39.
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Biographies Jon Spratley obtained his MEng in Mechanical Engineering in 2004, and is currently in his third year of studying for his PhD in the field of Microsystems and Nanotechnology at the University of Birmingham. His research interests are in neural and brain interfacing, and he is currently researching microstructure design for implantable systems. Dr. Ward obtained his physics degree from Imperial College in 1981 and subsequently was awarded his PhD from Warwick University for his EXAFS study on the passivation of stainless steels. In 1985 he joined RSRE Malvern where he worked for IR detectors, image processing, SOI electronics and MEMS before joining the University of Birmingham in 1999. He is currently senior lecture in microsystems engineering with research interest in noise, non-linear mechanics and microfluidics. Peter Hall is Professor of Communications Engineering, leader of the Antennas and Applied Electromagnetics Laboratory, and Head of the Devices and Systems Research Centre in the Department of Electronic, Electrical and Computer Engineering at The University of Birmingham. After graduating with a PhD in antenna measurements from Sheffield University, he spent 3 years with Marconi Space and Defence Systems, Stanmore working largely on a European Communications satellite project. He then joined The Royal Military College of Science as a Senior Research Scientist, progressing to Reader in Electromagnetics. He joined The University of Birmingham in 1994. He has researched extensively in the areas of microwave antennas and associated components and antenna measurements. Clive Thursfield started his career with an honours degree in electronics followed by a PhD in Biomedical Engineering from Strathclyde University where his research was associated with non-invasive cardiac monitoring. He is currently the Head of Research for the West Midland Rehabilitation Centre (WMRC) and for South Birmingham PCT. He is also the head of a newly created Department of Health Care Science (within WMRC) whose remit will be to ensure the use of scientific approaches and innovative and emerging technologies across the whole spectrum of rehabilitation. Clive is a Chartered Engineer and a Chartered Scientist.