Design and Characterization of a Small-Scale ... - IEEE Xplore

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IEEE/ASME TRANSACTIONS ON MECHATRONICS, VOL. 18, NO. 1, FEBRUARY 2013

Design and Characterization of a Small-Scale Magnetorheological Damper for Tremor Suppression David Case, Behzad Taheri, and Edmond Richer

Abstract—This paper explores the design methodology and effectiveness of small-scale magnetorheological dampers (MRDs) in applications that require variable damping. Previously, applications of MRD have been chiefly limited to vehicle shock absorbers and seismic vibration attenuators. There has been recent biomedical interest in active-damping technology, however, particularly in the field of rehabilitation robotics. The topic at hand is the feasibility of developing MRDs that would be functionally and cosmetically adequate for actuation of an upper limb tremor suppression orthosis. A Bingham plastic model is used to determine MRD’s functional characteristics, and experimental data are presented to validate the mathematical model. The feasibility of applying the developed small-scale MRDs to attenuation of tremorous motion is explored. Index Terms—Magnetic liquids, medical robotics, orthotics.

I. INTRODUCTION A. Tremor and Current Clinical Treatment REMOR is clinically described as a rhythmical, involuntary oscillatory movement of a body part produced by reciprocally innervated antagonist muscles and can be divided into two primary categories of movement disorders: resting and action tremor [1]. Resting tremor, most commonly associated with Parkinson’s disease, arises after a brief period of nonuse of the target muscle or muscle group. While not particularly debilitating in and of itself, resting tremor can be the cause of social embarrassment. Action tremor, in contrast, becomes apparent during muscle use. The tremor typically manifests at a frequency in the range of 3–12 Hz and can be particularly debilitating to fine motor skills, such as is needed in writing, shaving, eating, or playing an instrument. Many patients complain of social embarrassment, and some have been driven to career changes. Current treatments for various action tremors include a collection of prescription drugs and, in especially debilitating or nonresponsive cases, neurosurgery [2]. The medications typically prescribed for essential tremor may cause the patient to experience excessive drowsiness, nausea, ataxia, confusion, blurred vision, fatigue, and even muscle paralysis and hallucinations. Deep brain stimulation and stereotactic thalamotomy are mostly empirical surgical options that

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Manuscript received October 28, 2010; revised February 8, 2011; accepted April 2, 2011. Date of publication August 30, 2011; date of current version September 12, 2012. Recommended by Technical Editor W. J. Zhang. The authors are with the Biomedical Instrumentation and Robotics Laboratory, Southern Methodist University, Dallas, TX 75205 USA (e-mail: [email protected]; [email protected]; [email protected]). Color versions of one or more of the figures in this paper are available online at http://ieeexplore.ieee.org. Digital Object Identifier 10.1109/TMECH.2011.2151204

have been linked with permanent complications, paresthesia, dysarthria, speech impediment, and even stroke and hemiparesis [3], [4]. Essential tremor and Parkinson’s disease are degenerative conditions. Thus, while the administration of drugs, stereotactic thalamotomy, or thalamic deep brain stimulation is often initially effective in controlling tremor motion, none of these treatments guarantees a permanent solution [1], [2].

B. Tremor Orthoses Given the side effects and possible complications of current treatments for action and resting tremor, many researchers recognized the necessity of a less invasive alternative, the attenuation of tremor at the musculoskeletal level. In 1988, Sanes, LeWitt, and Mauritz reported on the use of a brushless dc motor to study the effects of adding viscous damping and inertia loads across the wrists of five patients experiencing various types of action tremors. Although the initial intention of the project was to shed light on the driving mechanisms of tremorous motion, the investigators reported that the application of viscous loads “suppressed the (local) tremor nearly linearly [5].” The Massachusetts Institute of Technology developed two orthoses for the purpose of evaluating and testing the effects of adding viscous and inertia loads more generally to the upper limbs of patients suffering from action tremor. The “controlled energy dissipation orthosis” (CEDO) was a wheelchair-mounted device, which functioned by applying resistive loads via magnetic particle brakes to a cuff attached to the patient’s wrist. The mechanism allowed the cuff 3 DOF in a horizontal plane. The “modulated energy dissipation manipulator” (MEDM) was the second generation of this device and less restrictive to general motion. The MEDM allowed the wrist cuff 6 DOF in 3-D space; however, it was also larger than the CEDO and nonportable. During evaluation of both devices, “the application of viscous damping loads was demonstrated to reduce tremor severity [6].” While no single set of damping parameters was observed to be most effective in attenuating tremor for all of the patients tested, the investigators concluded that “an individualized optimal level may exist which can maximize function [6].” The “viscous beam” orthosis, from the University of California Davis, was designed to attenuate tremor via viscous damping along a single degree of freedom in the forearm, specifically, flexion and extension of the wrist. The device proved successful in principle, although experimental values differed greatly from the calculated damping rates. The damping rate was not variable, and thus, the device’s degree of functional success was inconsistent between patients [7].

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CASE et al.: DESIGN AND CHARACTERIZATION OF A SMALL-SCALE MAGNETORHEOLOGICAL DAMPER FOR TREMOR SUPPRESSION

An orthosis acting on the same general principle was created in the course of the European “dynamically responsive intervention for tremor suppression” (DRIFTS) project, with the noted difference of employing a magnetorheological fluid (MRF) to alter the damping ratio between patients and potentially in real time. The design allowed tuning of the damping properties to optimize performance and was less restrictive to general motion. It allowed adduction/abduction and pronation/supination and added viscous damping to flexion/extension of the wrist [8], [9]. The “wearable orthosis for tremor assessment and suppression” (WOTAS) was the final product of the DRIFTS project. This device was designed to be minimally restrictive to natural movement and capable of monitoring and suppressing tremor, employing both active and passive strategies [10], [11]. The passive strategy employing a constant damping rate proved less effective than the active one. However, the investigators noted that “customization of viscosity or inertia added to the upper limb according to the biomechanical characteristics of each user should improve the efficiency of passive tremor suppression strategy [12].” The orthosis proved functionally successful in reducing the amplitude of the patients’ tremor by as much as 90%, though there were some complaints on the aesthetic drawbacks of the device, and the investigators stressed the necessity of further reducing its size and weight [13], [14]. The main challenge for the design and construction of an effective, least intrusive, portable tremor cancelling orthosis is the availability of compact, powerful, light, direct drive actuators and dampers. Thus, it is necessary to develop novel actuators/dampers that satisfy the power/force requirements for tremor cancellation, and have a sufficiently small profile and weight to be either hidden or easily ignored. In addition, to minimize interference with voluntary motion and avoid discomfort, they should exhibit the least amount of resistance force when back driven. None of the existing actuators combine all the desired characteristics. While electrical motors are easily powered by batteries and can be effectively controlled using relatively simple algorithms they cannot provide high-force, lowmechanical impedance actuation. Magnetorheological dampers (MRDs) excel at creating energy efficient variable damping forces but have relatively high mechanical impedance due to high viscosity of the fluid. In this paper, an MRD was designed using two mathematical models. The first one uses a planar approximation of the internal geometry and leads to a linear set of equations. The second model considers the real annular geometry of the damper resulting in a nonlinear model that is potentially more accurate. Two sets of experiments were conducted in order to validate the mathematical models and the applicability of MRDs for tremor suppression. II. MRFS An MRF consists of a suspension of microscopic magnetizable particles in a nonmagnetic carrier medium, usually water or some type of synthetic oil. In the absence of a magnetic field, the fluid behaves in a roughly Newtonian manner. When a magnetic field is present in the same space, the microscopic

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Fig. 1. Particle arrangement in the (a) absence and (b) presence of magnetic flux across the elements.

particles suspended in the fluid become oriented and form chains along the magnetic flux lines, changing the fluid’s rheology (see Fig. 1). Under these conditions, the properties of the fluid become nonisotropic. The presence of the particle chains causes the fluid to exhibit a yield stress if flow occurs perpendicular to the magnetic flux lines. Thus, when a magnetic field is applied, the MRFs behave similarly to a Bingham plastic. Since the observed yield stress is directly related to the intensity of the magnetic field, MRFs seem ideally suited for use in a low-power tunable damper. Variable resistance through the use of MRFs is a strategy currently being used commercially in vehicle shock absorbers and seismic vibration dampers for civic structures. In the field of human–machine interaction, particularly that of wearable robotics, high strength-to-weight ratio actuators are required to maximize assistive and rehabilitative potential [15]. Magnetorheological-based actuators can potentially achieve these high ratios and have the additional advantages of rapid response time and high fidelity control [16], [17]. These characteristics allow MRF dampers to be customized to the individual needs of a patient, as well as change the damping factor if the severity of the condition changes. III. GENERAL DAMPER DESIGN With the intent of minimizing the profile of a tremorsuppression orthosis, the proposed design incorporates a series of linear dampers/actuators acting roughly in parallel (in tension and compression) to the muscles of the forearm. A damper of the general design proposed here would be secured above the dorsal and radial surfaces of the forearm. The shafts would be connected to the hand by articulated linkages, thus, forming a direct-drive mechanism for applying torque at the transverse and dorsopalmar axes of the wrist. The palmar and ulnar surfaces would be unobstructed, allowing proximal movement and functionality at a desk. In an attempt to balance efficiency with ease of manufacturing, a piston/cylinder design was adopted for the dampers. A copper coil is wound about the piston head to create a magnetic field within the annular gap between the piston and interior cylinder surface (see Fig. 2). Both the cylinder wall and the piston head are machined from magnetically permeable material to close the magnetic circuit and direct the magnetic flux lines normal to the piston/cylinder gap. DRIFTS project investigators measured the biomechanical characteristics of tremorous movement at each joint in the arms of 33 patients, providing the necessary torque characteristics

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Fig. 2.


Cross-sectional diagram of a piston/cylinder MRD design.

TABLE I

Fig. 3.

Unit diagram for Bingham plastic flow between fixed parallel plates.

TORQUE MEASUREMENTS AT THE WRIST

The constructed damper weighs approximately 204 g. When run continuously with the maximum allowable current applied to its coil, the damper requires a 1.9 W power supply. for the actuators of a tremor-suppression orthosis [18]. The mean values of torque measured in the wrist are summarized in Table I. With a projected moment arm of 3 cm about the patient’s wrist joint, the required damper would need to be able to produce a resistance force of approximately 37 N. Given that the frequency of tremorous movement in the arm typically ranges between 3 and 12 Hz, the undamped velocity at the damper connection is projected at approximately 0.5–2.0 m/s. The primary variables that affect the resistance force in this piston/cylinder design are the shaft diameter Ds , piston diameter Dp , cylinder inside diameter Dc , active length L, and the fluid’s viscosity and yield strength (the latter being dependent upon the generated magnetic field and, thus, upon the dimensions of the coil and the applied current). The design required a hollow shaft to accommodate the coil leads that is both nonmagnetic and structurally rigid. Thus, the shaft diameter (4.76 mm) was based upon material availability. The cylinder inside the diameter was restricted early on in the design process to a maximum of 12.7 mm to keep the device’s profile reasonably small. In order to minimize resistance of the damper in passive mode, a low-viscosity MRF was selected for the project (LORD MRF-122EG, μ = 0.042 ± 0.020 Pa·s). The 0.635-mm piston/cylinder gap was selected in order to reduce the damping rate in passive mode, thus defining the piston diameter at 11.43 mm. The coil was restricted to the outer diameter of the piston. It was estimated that the coil could generate a magnetic flux density of 1.7 T at its core when the maximum admissible current of 0.54 A is applied. According to the manufacturer’s specifications, the required MRF yield strength of 15 kPa can be produced using a relatively small fraction of the maximum field intensity. Thus, the remaining design parameter (the active length of the piston) was determined through use of the mathematical model, requiring that the damper be able to produce at minimum a resistance force of 37 N at an operating velocity of 0.5 m/s, when maximum current is applied to the coil. By regulating the current, then, the resistance force of the damper can be varied in real time to suit the patient’s needs.

IV. MATHEMATICAL MODEL A. Linear Model For steady flow between parallel elements (i.e., the piston head and interior cylinder wall), Phillips provides a thorough description of ideal Bingham plastic behavior [19]. The derivation is illustrated with the help of Fig. 3, with the x-axis corresponding to the direction of flow and the z-axis normal to both surfaces. Summing forces on the unit ∂τ ∂P P δz + τ + δz δx − P + δx δz − τ δx = 0 (1) ∂z ∂x and can be reduced to dP dτ = . dx dz

(2)

For simplicity, the pressure gradient, −dP /dx, is denoted by P . Thus, dτ = −P dz

(3)

τ = C1 − P z.

(4)

By symmetry of the flow and shear profile, we have τ = 0 at z = (h/2). Therefore, C1 =

P h . 2

(5)

Let z = h1 at τ = τy , so that τy = C1 − P h1 = h1 =

P h − P h1 2

τy h − . 2 P

(6) (7)

Using hc to describe the “core flow” region that is farther than h1 from either of the parallel plates where τ < τy def

hc = h − 2h1 =

2τy . P

(8)


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Applying Bingham’s equations to the remaining “transient” region τ = τy + η

du dz

for τ > τy

(9)

τ = C1 − P z = (τy + P h1 ) − P z = τy + P (h1 − z) = τy + η

du dz

for 0 ≤ z ≤ h1 (10)

P (h1 − z) du = η dz u = C2 +

for 0 ≤ z ≤ h1 P (2h1 z − z 2 ) 2η

(11) Fig. 4.

for 0 ≤ z ≤ h1 .

A no-slip condition at the surface (u = 0 at z = 0) leads to Cz = 0, and u=

P (2h1 z − z 2 ) 2η

for 0 ≤ z ≤ h1 .

(13)

The core velocity evaluated at z = h1 = (h/2) − (τy /P ) is given by uc =

(P h − 2τy ) P h21 = . 2η 8P η

(14)

Using b to indicate the width of the elements, the core flow rate is signified by Qc = bhc uc = b(h − 2h1 )uc .

(15)

Integrating the velocity profile in the transient region, we can determine the flow rate there as h1 h1 P (2h1 z − z 2 ) dz u dz = 2b Qb = 2b 2η 0 0 bP 4 h31 2bP h31 3 = = buc h1 . (16) h1 − = η 3 3η 3 The total flow rate between the elements Q is the summation of Qc and Qb Qb + Qc 4 Q = = uc h1 + (h − 2h1 )uc . (17) b b 3 Substitution of the previously calculated values for uc and h1 leads to the following cubic equation, which governs the pressure gradient: τ 3 12 Q η τy y 3 2 + 3 + 4 = 0. (18) P − P b h3 h h Recalling the description of the core flow region, one recognizes that for any flow to occur, hc must be strictly less than h. When the terms are equated 2τy . (19) P Rearranging this equation yields the following minimum pressure gradient, below which there can be no flow: h=

Pm in

2τy . = h

Approximate velocity profile for a Bingham plastic in annular flow.

(12)

(20)

This minimum value determines which of the three solutions to the governing equation is relevant to the model. Thus, a readily solvable linear system is obtained to assist in the design of an MRD with specific desired characteristics. It should be noted that there are several assumptions employed in using this system to model the proposed piston-cylinder configuration. First, the formulae given here are for Bingham plastic flow between fixed elements, whereas, in a piston-cylinder device, flow occurs between the interior surface of the cylinder and the piston plunger, which moves relative to that surface. Second, these equations assume the plates to be flat and parallel, whereas the piston/cylinder device has a radial symmetry along its long axis. Finally, these calculations serve only to estimate the damping rate of the device; additional resistance forces (i.e., friction in the seals) are not accounted for but are ideally minimal. B. Nonlinear Model Kamath, Hurt, and Wereley describe a model of Bingham plastic flow that accounts for both the moving elements and the axial symmetry of a piston/cylinder device [20]. The Navier– Stokes equation, in cylindrical coordinates is used to model the force equilibrium ρ

τ ∂p ∂u ∂τ + + = . ∂t ∂r r ∂z

(21)

Neglecting inertia and assuming a constant pressure gradient over the characteristic length L dτ τ ΔP + = . dr r L

(22)

Just as before, the flow is divided into two “post-yield” or transient regions and one “core flow” region, producing a velocity profile in the gap as shown in Fig. 4. Applying Bingham’s description in the post-yield region adjacent to the piston, where shear stress exceeds the yield stress of the fluid du dr

(23)

dτ d2 u =η 2. dr dr

(24)

τ = τy + η

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The force balance becomes η

ΔP d2 u (τy + η(du/dr)) = . + dr2 r L

(25)

Integration leads to the velocity profile in region 1 u1 (r) =

ΔP 2 τy r − r + C1 ln(r) + C2 . 4ηL η

(26)

In the parallel-plate model there was symmetry across the core flow region, and it was safe to assume that the two post-yield regions were identical. This assumption cannot be made in the radial model, and the post-yield region adjacent to the cylinder, region 3, requires a slightly different description τ = −τy + η

Fig. 5.

du dr

dτ d2 u =η 2 dr dr

(28)

ΔP d2 u (−τy + η(du/dr) = + dr2 r L ΔP 2 τy r + r + C3 ln(r) + C4 . u3 (r) = 4ηL η η

(29) (30)

In region 2, velocity is uniform and defined by u2 = u1 (Rpi ) = u3 (Rpo ).

(31)

A volume balance can thus be established, equating flow through the annular gap with fluid displacement due to the motion of the piston profile Rp i 2 2 − v0 π(R1 − Rs ) = u1 2πr dr R1

R2

u3 2πr dr.

(32)

Rp o

The system has effectively three equations with four unknowns, ΔP, Rpi , Rpo , and u2 . To get the final equation necessary for a solution, the simplified Navier–Stokes equation is integrated dτ τ ΔP + = dr r L C5 ΔP r+ τ (r) = 2L r ΔP 2 r + C5 . r τ (r) = 2L

The shear stress is known at the limits of the core flow region, leading to the boundary conditions τ (Rpi ) = τy

τ (Rpo ) = −τy .

Substitution leads to the fourth necessary equation

The constants C1 , C2 , C3 , and C4 can be determined by applying the following boundary conditions, assuming no-slip conditions at the cylinder wall and the piston head (the velocity of which is −v0 ): du1 u1 (R1 ) = −v0 =0 dr r →R p i du3 = 0. u1 (R2 ) = 0 dr r →R p o

2 2 + u2 π(Rpo − Rpi )+

Experimental setup for damper characterization.

(27)

(33) (34) (35)

ΔP 2 R + C5 2L pi ΔP 2 R + C5 −Rpo τy = 2L po ΔP (Rpi + Rpo )τy = 2 − R2 ) 2L(Rpi po Rpi τy =

τy =

ΔP . 2L(Rpi − Rpo )

(36) (37) (38) (39)

One may note that the governing system of equations in this case is nonlinear, and, thus, more time consuming to solve. Once a set of acceptable dimensions was found with the parallel-plate model, it was evaluated with this nonlinear model as well. The calculated resistance force was found to differ by less than 0.8%. Thus, the linear model is considered sufficiently accurate for design purposes. V. EXPERIMENTS A. Model Validation The experimental setup built to evaluate the properties of the small-scale MRD includes a disk with inertial loads actuated by a brushless dc electrical motor and corresponding amplifier (EDC, Cambridge, MA). An optical encoder (ACCU-Coder 260 N-T-02-S-1000, Encoder Products Co., Sandpoint, ID) measures the angular position of the disk. The MRD is mounted in a rotating support and connected to the disk through a force transducer (MLP-50, Transducer Techniques, Temecula, CA). A multifunction RIONIPCI-7833R field-programmable gate array (FPGA) card allows real-time data acquisition from the encoder and force sensor and control of the motor. In order to mitigate the discrete nature of the optical encoder readout, the angular velocity and acceleration are estimated in real time using a custom-developed algorithm implemented in the internal memory of the FPGA card (see Fig. 5).


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Fig. 8. Experimental data for the resistance force of the damper at 3 Hz sinusoidal excitation. Fig. 6. Experimental values for the static resistance force of the damper at various applied currents.

Fig. 9. Experimental data for the resistance force of the damper at 6 Hz sinusoidal excitation.

Note that the same peak velocity and applied current produces approximately the same resistance forces at both frequencies, as expected. Fig. 7.

Yield force versus applied current.

B. Tremor Reduction Application

Static properties of the damper were measured by applying a slowly increasing linear force with various currents applied to the coil and measuring the position change to determine the point of yield as shown in Fig. 6. The linear mathematical model was used to determine the theoretical yield force under the experimental conditions. These theoretical values were given a bias of 3 N to account for friction in the seals and a gain of 0.10 to account for losses in the magnetic field. Good agreement is observed between measured and theoretical values for several applied currents as shown in Fig. 7. In order to partially characterize the dynamic properties of the damper, excitation frequencies of 3 and 6 Hz were imposed with a peak velocity of ∼0.2 m/s. Figs. 8 and 9 show the resistance force of the damper measured with the following configurations: no fluid in the chamber, MRF in the chamber (no current through the coil), and MRF in the chamber (20% max current through the coil).

In order to estimate the functional potential of the smallscale MRD in the construction of a tremor reduction orthosis, an additional set of experiments was performed. The voltage applied to the electric motor consisted of the superposition of two sinusoidal waves of two different frequencies, the lower frequency wave representing voluntary motion, and the higher frequency wave representing the involuntary tremor. Several cycles of motion were examined with an incremental increase of current through the coil. The experimental results and a fast Fourier transform of each dataset are shown in Fig. 10 in order to quantify the signal spectral composition. With no current applied to the MRD, the damping force is strictly a result of friction in the seals, the inherent viscosity of the fluid, and the damper’s inertial properties. With the application of current to the coil, one can note an immediate decrease in the amplitude of motion for both sinusoids and an increase of resistive force. As current is increased, these effects become more pronounced. At 200 mA, both voluntary and involuntary motion has been almost completely eliminated. At this point, there is effectively no fluid flow in the annular gap of the damper.

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Fig. 10. Response of the damper to the dual frequency excitation with an applied current of (a) 0, (c) 100, and (e) 200 mA and spectral composition of the response with an applied current of (b) 0, (d) 100, and (f) 200 mA.

VI. DISCUSSION AND CONCLUSION A small-scale MRD was designed and constructed based on a linear mathematical model of MRF flow under magnetic field. The design parameters obtained using the linear model were verified using a potentially more accurate nonlinear model that considers radial symmetry and relative motion of the internal surfaces of the damper. No significant differences where noted in the estimated resistance force using the two models. Experimental validation of the models was conducted. As expected, the modification of the magnetic field produced by the coil is shown to significantly alter the resistance force of the damper. Good agreement was observed between the values predicted by

the theory and the measured force when the friction and losses in the coil’s magnetic field were considered. Nevertheless, the relatively high losses in the magnetic field at the piston/cylinder gap shown experimentally emphasize the need to optimize the coil geometry. In addition, experiments designed to evaluate the MRD’s performance in a tremor reduction orthosis were performed. The damper attenuated both the voluntary and involuntary motions to a similar degree. In order to restore the voluntary motion, the amplitude of the low frequency signal applied to the electrical motor had to be doubled. This resulted in a voluntary motion with no significant tremorous disturbance. In orthotic


applications, this effect would translate into a significant resistance force felt by the wearer, but it would potentially improve the ability to perform fine motor skill tasks. The mathematical models presented in this paper assumed fully developed flow within the damper, and fluid inertia was not taken into account. These models are, therefore, suitable to predict the operational range of the damper, but they are inadequate to accurately predict behavior in oscillatory flow, which would be seen in application to tremor reduction. In order to develop an effective control strategy for the attenuation of tremor with MRDs, further modeling (e.g., taking fluid inertia and response time into account) is necessary. REFERENCES [1] G. Deuschl, P. Bain, and M. Brin, “Consensus statement of the movement disorder society on tremor,” Movement Disorders, vol. 13, no. 3, pp. 2–23, 1998. [2] D. Sirisena and D. R. Williams, “My hands shake: Classification and treatment of tremor,” Austral. Fam. Phys., vol. 38, no. 9, pp. 678–683, Sep. 2009. [3] Y. Katayama, T. Kano, K. Kobayashi, H. Oshima, C. Fukaya, and T. Yamamoto, “Difference in surgical strategies between thalamotomy and thalamic deep brain stimulation for tremor control,” J. Neurol., vol. 252, no. 4, pp. 17–22, 2005. [4] E. D. Flora, C. L. Perera, A. L. Cameron, and G. J. Maddern, “Deep brain stimulation for essential tremor: A systematic review,” Movement Disorders, vol. 25, no. 11, pp. 1550–1559, 2010. [5] J. N. Sanes, P. A. LeWitt, and K.-H. Mauritz, “Visual and mechanical control of postural and kinetic tremor in cerebellar system disorders,” J. Neurol., Neurosurg., Psychiatry, vol. 51, pp. 934–943, 1988. [6] M. L. Aisen, A. Arnold, I. Baiges, S. Maxwell, and M. Rosen, “The effect of mechanical damping loads on disabling action tremor,” Neurology, vol. 43, pp. 1346–1350, Jul. 1993. [7] J. Kotovsky and M. J. Rosen, “A wearable tremor-suppression orthosis,” J. Rehabil. Res. Develop., vol. 35, pp. 373–387, Oct. 1998. [8] R. C. V. Loureiro, J. M. Belda-Lois, E. R. Lima, J. L. Pons, J. J. SanchezLacuesta, and W. S. Harwin, “Upper limb tremor suppression in ADL via an orthosis incoporating a controllable double viscous beam actuator,” in Proc. IEEE 9th Int. Conf. Rehabil. Robot., Chicago, IL, Jun./Jul. 2005, pp. 119–122. [9] J.-M. Belda-Lois, A. Page, J.-M. Baydal-Bertomeu Rakel Poveda, and R. Barbera, Rehabilitation Robotics: Biomechanical Constraints in the Design of Robotic Systems for Tremor Suppression. Vienna, Austria: I-Tech, 2007. [10] M. Manto, M. Topping, M. Soede, J. Sanchez-Lacuesta, W. Harwin, J. Pons, J. Williams, S. Skaarup, and L. Normie, “Dynamically responsive intervention for tremor suppression,” Rehabil. Robot., vol. 22, pp. 120– 132, May/Jun. 2003. [11] E. Rocon, M. Manto, J. Pons, S. Camut, and J. M. Belda, “Mechanical suppression of essential tremor,” Cerebellum, vol. 6, pp. 73–78, 2007. [12] E. Rocon, J. M. Belda Louis, A. F. Ruiz, M. Manto, J. C. Moreno, and J. L. Pons, “Design and validation of a rehabilitation robotic exoskeleton for tremor assessment and suppression,” IEEE Trans. Neural Syst. Rehabil. Eng., vol. 15, no. 3, pp. 367–378, Sep. 2007. [13] E. Rocon, A. F. Ruiz, and J. L. Pons, “Rehabilitation robotics: A wearable exo-skeleton for tremor assessment and suppression,” in Proc. 2005 IEEE Int. Conf. Robot. Autom., Barcelona, Spain, Apr. 2005, pp. 2271–2276. [14] E. Rocon, M. Manto, J. Pons, J. M. Belda, and S. Camut, “Evaluation of a wearable orthosis and an associaed algorithm for tremor suppression,” Physiol. Meas., vol. 28, no. 4, pp. 415–425, Apr. 2007. [15] J. Rosen, S. Burns, and J. C. Perry, “Upper-limb powered exoskeleton design,” IEEE/ASME Trans. Mechatronics, vol. 12, no. 4, pp. 408–417, Aug. 2007. [16] H. B. Gurocak and J. Blake, “Haptic glove with MR brakes for virtual reality,” IEEE/ASME Trans. Mechatronics, vol. 14, no. 5, pp. 606–615, Oct. 2009. [17] A. S. Shafer and M. R. Kermani, “On the feasibility and suitability of MR fluid clutches in human-friendly manipulators,” IEEE/ASME Trans. Mechatronics, [Online]. Available: http://ieeexplore.ieee.org, DOI: 10.1109/TMECH.2010.2074210.

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[18] J. M. Belda-Lois, E. Rocon, J. J. Sanchez-Lacuesta, A. F. Ruiz, and J. L. Pons, “Estimation of the biomechanical characteristics of tremorous movements based on gyroscopes,” in Assistive Technol.: From Virtuality Reality, vol. 16, A. Pruski and H. Knops, Eds. Amsterdam, The Netherlands: IOS Press, 2005, pp. 138–142. [19] R. W. Phillips, “Engineering applications of fluids with a variable yield stress,” Ph.D. dissertation, Dept. Mech. Eng., Univ. California, Berkeley, 1969.. [20] G. M. Kamath, M. K. Hurt, and N. M. Wereley, “Analysis and testin of bingham plastic behavior in semi-active electrorheological fluid dampers,” Smart Mater. Struct., vol. 5, pp. 576–590, 1996.

David Case received the B.S. degree in mathematics from the University of Dallas, Irving, TX, in 2007. He is currently working toward the M.S. degree in the Biomedical Instrumentation and Robotics Laboratory, Southern Methodist University, Dallas, TX. He is currently a Research Assistant in the Biomedical Instrumentation and Robotics Laboratory, Southern Methodist University. In 2008–2009, he assisted studies in multiple robot interaction at the Automation and Robotics Research Institute, The University of Texas at Arlington. His current research interests include rehabilitation robotics and novel actuator design with smart materials.

Behzad Taheri received the B.S. degree from Bu Ali Sina University, Hamedan, Iran, and the M.S. degree from Sharif University of Technology, Tehran, Iran, both in mechanical engineering. He is currently working toward the Ph.D. degree in mechanical engineering in the Biomedical Instrumentation and Robotics Laboratory, Southern Methodist University, Dallas, TX. He is currently a Research Assistant in the Biomedical Instrumentation and Robotics Laboratory, Southern Methodist University. His current research interests include dynamics of human gait, human arm tremor control, rehabilitation, nonlinear control, and robust control.

Edmond Richer received the M.S. degree in mechanical engineering from the University of Craiova, Craiova, Romania, in 1988 and the Ph.D. degree in dynamic systems and control from the Lyle School of Engineering, Southern Methodist University, Dallas, TX, in 1998. He currently directs the Biomedical Instrumentation and Robotics Laboratory, Southern Methodist University, where he is an Associate Professor in the Bobby B. Lyle School of Engineering. His current research interests include advanced dynamics and control, image-guided and haptic-assisted robotic systems, and ultrasound applications in bone quality assessment and supra resolution imaging. Dr. Richer is currently a member of the American Society of Mechanical Engineers.