IEEE TRANSACTIONS ON ROBOTICS, VOL. 24, NO. 1, FEBRUARY 2008 159
A Gas-Actuated Anthropomorphic Prosthesisfor Transhumeral Amputees
Kevin B. Fite, Member, IEEE, Thomas J. Withrow, Member, IEEE, Xiangrong Shen,Keith W. Wait, Member, IEEE, Jason E. Mitchell, and Michael Goldfarb, Member, IEEE
Abstract—This paper presents the design of a gas-actuated an-thropomorphic arm prosthesis with 21 degrees of freedom and nineindependent actuators. The prosthesis utilizes the monopropellanthydrogen peroxide as a gas generator to power nine pneumatic typeactuators. Of the nine independent actuators, one provides direct-drive actuation of the elbow, three provide direct-drive actuationfor the wrist, and the remaining five actuate an underactuated 17degree of freedom hand. This paper describes the design of theprosthesis, including the design of small-scale high-performanceservovalves, which enable the implementation of the monopropel-lant concept in a transhumeral prosthesis. Experimental resultsare given characterizing both the servovalve performance and theforce and/or motion control of various joints under closed-loopcontrol.
Index Terms—Monopropellant, myoelectric, pneumatic,prosthesis, transhumeral.
I. INTRODUCTION
R ECENT advances in neural and mechanical interface tech-nology bring to the horizon the potential to significantly
enhance the capability of upper extremity prostheses. Specifi-cally, the recent approaches to neural interfacing (see, for ex-ample, [1]–[14]) have the potential to provide a highly func-tional prosthesis with a high number of control inputs, and re-cent work in mechanical interfaces, such as the osseointegrationtechniques described in [15]–[18], provides the potential to in-tegrate the structure of such a prosthesis directly into the user’sskeleton. Leveraging such advances requires the development ofhighly functional anthropomorphic upper extremity prostheses
Manuscript received January 31, 2007; revised xxx. This paper wasrecommended for publication by Associate Editor B. Hannaford and Editor F.Park upon evaluation of the reviewers’ comments. This work was supportedby the Defense Advanced Research Projects Agency (DARPA) under ContractN66001-06-C-8005.
K. B. Fite is with the Department of Mechanical and AeronauticalEngineering, Clarkson University, Potsdam, NY 13699-5725 USA (e-mail:[email protected]).
This paper has supplementary downloadable multimedia material availableat http://ieeexplore.ieee.org, provided by the author. This material includes onevideo (IEEETRO arm video.wmv) demonstrating the Vanderbilt University 21DOF (9 actuator) arm. The arm is commanded by a master exoskeleton. Thevideo (can be played with Windows) has three segments, the first demonstratingthe arm running on a cold gas supply, the second demonstrates the audible soundcharacteristics, and the third shows the arm running on the monopropellant hotgas supply. The size is 35 MB. Contact [email protected] forfurther questions about this work.
Color versions of one or more of the figures in this paper are available onlineat http://ieeexplore.ieee.org.
Digital Object Identifier 10.1109/TRO.2007.914845
with significantly more powered degrees of freedom than thosecurrently available. Among the many challenges that exist inthe development of an anthropomorphic arm is that of provid-ing a high number of actuated degrees of freedom, each capableof significant force and/or power, in a highly confined space.As such, a critical component of any viable approach to an an-thropomorphic arm is the incorporation of actuators that delivercomparable power density, force density, strain capability, andbandwidth to human skeletal muscle, and, importantly, mustembody a similar form factor. In fact, since some amputeesmay have a long residual limb, the accommodation of theseamputees will require a design in which the elbow actuator islocated within or distal to the elbow, in which case compara-ble performance to a human will actually require greater powerdensity than human skeletal muscle.
State-of-the-art electromagnetic motors have a reasonablepower density and a good bandwidth, but they exhibit a lowtorque density (relatively to human joint actuation), and do notprovide an appropriate form factor for high-density actuation inan anthropomorphic form. Due to their low torque density, anelectromagnetic motor requires a significant transmission ratio,typically embodied by a gearhead. Use of a gearhead reducessignificantly the power density of the actuator due to both areduction in efficiency and a corresponding increase in actuatorweight (i.e., the gearhead both decreases the numerator and in-creases the denominator of the power density). Specifically, thepower density of mammalian skeletal muscle is approximately150–250 W/kg [19], [20]. A state-of-the-art rare-earth-magnet-brushed motor with harmonic drive gearhead has a power den-sity of approximately 50 W/kg, and a state-of-the-art brushlessmotor with a planetary gearhead of about 100 W/kg (not includ-ing power electronics, the mass of which can be significant).As such, on average, the motor/gearhead combinations are ap-proximately three times less power dense than skeletal muscle.The combined drawbacks of low power density and awkwardform factor introduce significant challenges into the develop-ment of an electromagnetic-motor-actuated anthropomorphicarm with near-human capability. This paper describes an al-ternative means of developing an anthropomorphic arm basedon a gas-actuation approach. The proposed approach providesan actuator power density, force density, bandwidth, and formfactor conducive to the development of an anthropomorphicarm.
II. DESCRIPTION OF THE ACTUATION APPROACH
Gas-type actuators provide significantly better gravimetricand volumetric power density relative to electromagnetic-type
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actuators [21], and thus, provide an attractive possibility forthe development of an anthropomorphic prosthesis. In addi-tion to their high power density, gas-type actuators have sev-eral other properties, which make them particularly suitablefor use in an anthropomorphic prosthesis. Among these, gas-type actuators exhibit a low-output impedance relative to non-backdrivable gearheads, which enhances their ability to performconstrained interaction tasks [22] and decreases their sensitiv-ity to shock loading. Gas actuators do not require power con-sumption during isometric contraction, which is an importantcharacteristic when operating within a gravitational environ-ment. Finally, gas-type actuators have a form factor, whichis conducive to a high degree of freedom anthropomorphicarm.
One of the most significant issues when considering the useof gas-type actuators is determining how to power them. Anycompressor-based approach would have such a low power den-sity that it would completely offset any advantages of usinggas-type actuators. One possibility is to use the pressure-basedphase change characteristics of liquid carbon dioxide (CO2)to effectively store gas in a liquid form. In fact, gas-actuated,CO2-powered upper extremity prostheses were developed byseveral groups in the 1960s and 1970s [23]–[28]. Marquardtdescribes his experience in fitting more than 350 patients withCO2-powered pneumatic prostheses, and credits Hafner withinitiating the concept in 1948 [23]. Marquardt additionally re-ports that the CO2 cartridges would last a full day when theprosthesis was in “full time” use.
Despite the apparent successes of these prostheses, two majorissues appear to have impeded the long-term practicality of suchsystems. First, the CO2 cartridges were large and heavy relativeto an upper extremity prosthesis, with a typical capacity of 350mL and a tank mass of approximately 0.9 kg [29]. Second,the previous systems lacked feasible methods for stable andaccurate servo control, due largely to the absence of small-scale high-performance servovalves and, to a lesser extent, tothe absence of appropriate embedded control platforms (i.e.,microcontrollers).
The work presented in this paper addresses both of theseshortcomings (i.e., inadequate energy density and the absence ofsmall-scale high-performance servovalves), and as such makesfeasible a self-contained gas-powered anthropomorphic upperextremity prosthesis. Specifically, the energy density issue hasbeen addressed by incorporating the use of a liquid monopro-pellant as a gas generator in place of a CO2 cartridge, whichprovides roughly an order of magnitude improvement in energydensity. Additionally, the authors have developed small-scalehigh-performance servovalves, which enable accurate force andmotion control of the gas actuators in the weight and volumeconstraints of an anthropomorphic arm.
Regarding the energy density issue, the authors have con-ducted considerable work in the use of the monopropellanthydrogen peroxide for the purpose of powering gas actua-tors [30]–[32]. A monopropellant is a substance that reactsor decomposes exothermically when in contact with a cata-lyst. Since these reactions break molecular bonds, they can pro-vide a significantly greater energy density than a phase change,
TABLE IENERGY DENSITY OF GAS GENERATORS
Fig. 1. Schematic diagram of proposed actuator configuration.
which does not alter the molecular structure of the substance.Table I lists the gravimetric and volumetric energy density ofvarious concentrations of the liquid monopropellant hydrogenperoxide relative to liquid CO2 assuming a constant pressureprocess (isobaric conditions), where the energy density is thevolume of compressible gas produced at a given pressure perunit mass or volume of liquid. Specifically, the energy densityis given by RT, where R is the particular gas constant and T isthe gas temperature, which is assumed to be room temperaturefor the CO2 (an upper bound, which will only occur for veryslow discharge) and adiabatic decomposition temperature forthe H2O2 (which should be the case, regardless of dischargerate). Given that the temperature of the reaction increases withincreasing propellant concentration, feasible concentrations fora prosthesis are most likely limited to 70%–80%, which cor-responds to adiabatic decomposition temperatures of 232 ◦Cand 487 ◦C, respectively [33]. Based on the improvement involumetric energy density, a tank of 70%–80% H2O2 with acomparable energy content to CO2 would be five–seven timessmaller, with a corresponding decrease in the tank weight. Assuch, the 350 mL, 900 g tank described in [29] could be re-placed with a 70 mL, 180 g tank of 70% H2O2 , thus providingthe equivalent energy content in a small, lightweight package,relative to the space and mass constraints of an upper extremityprosthesis.
Fig. 1 depicts schematically the embodiment of themonopropellant-powered actuation approach as utilized in theprosthesis. In this embodiment, a small (8 g) CO2 cartridge isused to pressurize the propellant cartridge (i.e., is used as afuel pump), which is, in turn, connected to series of servovalvesthrough a catalyst pack. When the impedance downstream of thecatalyst pack is infinite (i.e., all gas valves are closed), the sys-tem will not draw any flow from the fuel cartridge (technically,a propellant cartridge). When the impedance downstream of thecatalyst pack is lowered (i.e., by opening one or several valves),
FITE et al.: GAS-ACTUATED ANTHROPOMORPHIC PROSTHESIS FOR TRANSHUMERAL AMPUTEES 161
pressure in the cartridge forces propellant through the catalyst,which effectively increases the flow rate from a signal level to apower level (i.e., through its exothermic release of heat and re-sulting gaseous expansion). As such, pressurized gas is suppliedon demand and used to power a set of gas actuators, whereinthe gas flow for each actuator is controlled via a respectivefour-way servovalve. The implementation of such an approachis, thus, highly dependent on the existence of scale-appropriateservovalves that have appropriate gas flow control character-istics, and that can accommodate the elevated gas tempera-tures associated with the reaction products of the monopropel-lant. Since such valves are not commercially available, the au-thors developed the appropriate servovalves, and describe themherein.
The proposed actuation approach enables actuation powerdensities comparable to skeletal muscle, and therefore, enablesthe development of a high degree of freedom anthropomorphicarm. Use of a liquid fuel, however, brings with it some dis-advantages, most notably: 1) production of exhaust products;2) some degree of audible noise associated with the exhaust;3) elevated internal reaction temperatures; and 4) increased dif-ficulty in packaging relative to a solid energy source such as abattery. Due to the inherently safe nature of the propellant (inconcentrations less than 80%), in combination with proper de-sign attention, these issues can be reduced to acceptable levels.Specifically, with regard to the safety of the propellant, hydro-gen peroxide at the proposed concentration levels (70%–80%)is completely nonflammable, insensitive to shock, has no possi-bility of detonation, and has completely safe reaction products(oxygen and steam). Despite this, the propellant is a strong ox-idizer, and, as such, is best distributed in sealed cartridges forconsumer distribution, much like CO2 or propane. With regardto the exhaust temperature, though the adiabatic reaction tem-perature (i.e., the temperature prior to any extraction of work orheat loss) at the catalyst pack is relatively high (232 ◦C for 70%peroxide), the gas temperature is quickly cooled as it expandsthrough the actuator and performs work. Though the exact ex-haust temperature depends on the motion and load, all priorexperimental experiences (with many types of loads and mo-tions) indicate that the exhaust temperature is comfortable forprolonged human exposure (i.e., one can hold one’s hand 1 infrom the exhaust outlet continuously during operation). Aver-age cylinder temperatures during typical operation have beenmeasured in bench-top experiments at 85 ◦C (185◦ F), whichis sufficiently low, such that a user could be adequately pro-tected from the hot surfaces with standard polyurethane foaminsulation. With regard to the audible output, the sound levelof the prosthesis prototype (as demonstrated in the accompa-nying video available at http://ieeexplore.ieee.org) measures ator below 50 dB from 1 m away, which, in most applications, isconsidered ambient.
This paper describes the design of the monopropellant-powered prosthesis, including the design of the small-scaleservovalves, and presents data indicating the closed-loopperformance of various degrees of freedom of the de-vice. An accompanying video demonstrates the functioningprosthesis.
Fig. 2. Four-way servovalve shown with an AA-size battery.
III. SERVOVALVE DESIGN
A major component necessary for the realization of the liquid-fueled arm prosthesis is a small-scale four-way servovalve thatcan accommodate elevated gas temperatures. As previouslymentioned, such a valve was not commercially available, andtherefore, was developed by the authors. The resulting servo-valve, which is shown next to an AA-size battery in Fig. 2, canaccommodate gas pressures of 2.1 MPa (300 lbf/in2) and gastemperatures of 232 C (450 F); has a flow factor of Kv = 0.20(flow coefficient of Cv = 0.25); has a bandwidth of approxi-mately 35 Hz; and has a mass (as shown in Fig. 2) of 28 g.A major design distinction between the valve shown in Fig. 2and typical direct-acting servovalve designs (e.g., Enfield Tech-nologies models LS-V05 and LS-V15, or Festo model 154200MPYE-5-M5-010-B) is that the design described here is rotaryrather than linear in nature. That is, the direction and extent ofgas flow is controlled by rotating the spool inside the sleeve,as described in [34], rather than by translating the spool withinthe sleeve, as is more typical. A rotary approach enables theuse of a gearhead, which better matches the power output ofthe servomotor to the valve-spool load, and enables the use ofa compact magnetic encoder for purposes of valve spool po-sition feedback control. An exploded view of the valve designis shown in Fig. 3. The design incorporates a precision groundspool, which is 3 mm in diameter, matched to a sleeve with adiametral clearance of 1 µm (XX gage quality) actuated by adc motor/encoder package (Micromo model 1319T006SR-IE2-400). The dc motor/encoder drives the valve spool through a9:1 custom cable-drive gearhead, which is shown in Fig. 4. Thegearhead consists of two 3:1 stages, whereby the motion of theshaft and roller of each stage are coupled by a pretensioned125-µm (0.005-in) diameter nylon monofilament cable (as seenin Fig. 4). The gearhead serves the dual purpose of a zero-backlash compact gear reduction and thermal isolation, the latterdue to the use of polyetheretherketone (PEEK) components.
The measured flow characteristics of the valve as a functionof the spool angle are shown in Fig. 5, where a positive massflow rate corresponds to flow between supply and the “A” port,
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Fig. 3. Exploded view of rotary servovalve.
Fig. 4. Two-stage cable-drive gearhead, shown with a U.S. penny in thebackground.
while a negative mass flow rate indicates flow between supplyand the “B” port. Note that the valve provides a smooth andfairly linear relationship, with a deadband of approximately 6◦
around the zero position. Note that such deadbands are commonin servovalves, since overlap between ports is required to ensurea positive shut-off at the zero position. The measured frequencyresponse of the valve (for a ±15◦ command) is shown in Fig. 6,indicating a –3 dB bandwidth of approximately 35 Hz, and arepresentative sinusoidal tracking response at 20 Hz is shown inFig. 7.
IV. FOREARM DESIGN
The actuators of the arm were sized to provide each jointwith the appropriate maximum torque and range of motion. Thedesired specifications for maximum torque and range of mo-tion for the elbow and each wrist degree of freedom are given inTable II, which were derived largely based on the output charac-teristics of a typical human arm, as characterized by [35]–[37].
Fig. 5. Measured flow characteristics of valve as a function of spool angle.
Fig. 6. Measured frequency response of valve spool for a ±15◦ command.The solid line shows a second-order fit of the measured data. Note that differ-ences between the fit and data are likely due to spool/sleeve friction and motorsaturation (torque and speed, due to current and voltage saturations).
TABLE IIDESIRED FOREARM ACTUATION SPECIFICATIONS
Also given in Table II is the total required energy that must bedelivered by each respective actuator, and the total required ac-tuator volume at an (maximum) operating pressure of 2.1 MPa(300 lbf/in2).
Based on the requisite displaced volume for each degree offreedom, actuator dimensions were determined, largely basedon the constraint that the prosthesis fit within the anatomical en-velope of the human arm. Table III summarizes the dimensionsof each actuator corresponding to the requirements of Table II,and provides an indication of the additional actuator energy
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TABLE IIIFOREARM ACTUATOR SPECIFICATIONS
Fig. 7. Sinusoidal spool tracking at 20 Hz for a±15◦ command, correspondingto the 20 Hz data points shown in Fig. 6.
available to overcome frictional losses in the device (e.g., jointand cylinder seal friction). Specifically, the ratio of actual dis-placed volume to the required displaced volume indicates therelative amount of additional energy available to overcome fric-tional losses (i.e., a ratio of 1.2 indicates a 20% “overdesign” ofavailable actuator energy).
Fig. 8 depicts a solid model of the elbow joint designed forthe anthropomorphic arm. The elbow joint consists of a revolutejoint actuated by a heavily modified flat cylinder (Bimba modelFO-093-3V) with a force sensor (Sensotec model AL311 CN)integrated in series with the cylinder piston rod. A custom fab-ricated four-way servovalve, as previously described and shownin Fig. 2, connects to the cylinder ports in order to control theflow of gas to and from the opposing cylinder ports. The revo-lute joint, shown in the exploded view of Fig. 9, is composed ofan inner link sandwiched between split outer links and incorpo-rates a potentiometer (ALPS model RDC503013 A) for angularposition sensing. The potentiometer is housed within the medialouter link and interfaces with the joint through a shaft pressedinto the inner link. A pair of flanged bearings (GGB modelBB0604DU) provides for low-friction rotation of the revolutejoint. The resulting design provides a range of motion of 105◦
with hard stops integrated within the split outer links of therevolute joint.
Pronation/supination of the wrist is actuated using a vari-ant of a leadscrew assembly, the design of which is shown in
Fig. 8. Solid model of elbow joint.
Fig. 9. Exploded view of elbow joint.
Fig. 10. Solid model of wrist pronation/supination.
Fig. 10. Specifically, the design consists of two concentric tubes,one inside the other. The inside tube contains a pair of straightslots, while the outer tube contains a pair of helical slots. Theactuation cylinder moves a pin (PEEK crosspiece in the fig-ure) through the slots, which rotates the outer tube relative tothe inner tube. The inner tube (the distal forearm tube) is sup-ported by a pair of flanged bearings (GGB models BB1612DUand BB1817DU) to allow rotation within the outer tube (theproximal forearm tube, shown partially transparent), while also
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Fig. 11. Solid model of wrist flexion/extension and abduction/adduction.
providing for thrust load bearing. The proximal forearm houses apotentiometer (ALPS model RDC503013A) to measure angularrotation of the joint. The distal forearm houses the double-actingcylinder (Bimba model 021.5-DXPV), which actuates the cross-piece through a force sensor (Measurement Specialties modelELFS-T4E-100L). The total angular range of motion of the jointis 95◦, as determined by the pitch of the helical slots.
The wrist flexion/extension and abduction/adduction degreesof freedom are shown in the solid model of Fig. 11, andconsist of a pair of cantilevered revolute joints actuated by apair of cylinders (Bimba models 022-DXPV and 021-DXPV).Each cylinder includes a load cell (Measurement Specialtiesmodel ELFS-T4E-100L) in series with its piston rod for forcesensing in each joint. The flexion/extension cylinder has simplepinned joints at each end, but due to coupling in the cantileveredjoint pair, the abduction/adduction cylinder must have sphericaljoints at each end. The cantilevered joint pair integrates poten-tiometers (ALPS model RDC503013A) and two pairs of flangedbronze bushings, as shown in the exploded view of Fig. 12. Theresulting design provides a range of motion of 105◦ for wristflexion/extension and 40◦ for wrist abduction/adduction.
V. HAND DESIGN
The prosthetic hand has 17 joints, which are actuated by fiveindependent actuators, and thus, is highly underactuated. Notethat the native hand is also underactuated, but to a lesser extentthan the design described here. The five actuators are allotted tothe 17 degrees of freedom in the hand, as described in Table IV.In all the cases, the underactuation is governed by momentisotropy, rather than by kinematic constraints. In other words,the hand will reach a configurational equilibrium when all jointmoments are (essentially) equal, excepting tendon friction andslight nonlinearities in the relationship between tendon forceand joint moment as a function of joint angle. Note that this isachieved by a combination of having the tendon span multiplejoints, and by using pulley differentials to split the force of the
Fig. 12. Exploded view of cantilever wrist joint pair.
TABLE IVDISTRIBUTION OF HAND ACTUATION
actuator output equally into two tendons, as briefly described inTable IV.
All joints in the hand are fully compliant (i.e., they lack shaftsand bearings). As shown in Fig. 13, each joint consists of a pairof oppositely wound torsional springs, which are supplementedwith two sets of inserts that together diminish compliance inoff-axis directions, and thus, enable a better approximation ofrevolute joint kinematics. Specifically, a split tube insert signifi-cantly increases axial mode stiffness, as illustrated in Fig. 14(a),and a ring insert significantly increases frontal plane bendingstiffness, as illustrated in Fig. 14(b).
Use of compliant joints in the hand serves several importantpurposes. First, and perhaps most obviously, the complianceprovides a return force, which simplifies tendon actuation ofthe joints, since active extension of the joints via an extensivetendon is not necessary. Second, the compliant joints map jointmotion to tendon force in free space, and thus: 1) eliminate theneed for position sensing in the hand and 2) eliminate the need toswitch between motion control and force control. Specifically,the hand frequently engages in both motion control (e.g., whengesturing or reaching) and force control (e.g., when grasping or
FITE et al.: GAS-ACTUATED ANTHROPOMORPHIC PROSTHESIS FOR TRANSHUMERAL AMPUTEES 165
Fig. 13. Finger assembly shown assembled and exploded.
Fig. 14. (a) Split tube insert diminishes longitudinal mode of deformation.(b) Ring insert diminishes frontal plane bending.
squeezing). The compliant joints offer a reliable, passive, instan-taneous means of switching between motion and force controlwithout any sensors or switches. Specifically, the tendons arealways exclusively under force control. When the fingers arenot in contact with other objects or fingers, the compliance atthe joints maps force to position, so the fingers are, in fact,under position control. When a finger comes in contact with arigid object, the mapping changes instantaneously from positionto force control. Compliant joints also provide insensitivity toshock, which is an important characteristic in prosthetic devices,and enables limited deformation along other axes such as fingerabduction.
A prototype of the hand, fabricated from acrylonitrile butadi-ene styrene (ABS) plastic and mounted on a test stand in threepositions is shown in Fig. 15. Note that seven cables actuate thehand, since two of the five-actuator inputs are split by pulleydifferentials at the actuator.
Fig. 15. (a) ABS prototype of the 17 degrees of freedom hand fully ex-tended with the use of torsional springs. (b) Demonstration of a spherical grasp.(c) Demonstration of a cylindrical tool grip.
TABLE VDESIRED HAND ACTUATION SPECIFICATIONS
TABLE VIHAND ACTUATOR SPECIFICATIONS
Actuators for the hand were sized in a similar manner tothose selected for forearm actuation, although they were basedon desired tendon force and excursion rather than their angularcounterparts. The desired specifications for maximum tendonforce and range of motion (i.e., tendon excursion) for the handactuators is given in Table V, based on typical human capabili-ties, as given in [35]–[37]. Also, given in Table V is the energythat must be delivered by each respective actuator to achievethe desired tendon specifications and the additional energy re-quired to overcome the joint compliance. Finally, the total re-quired actuator volume at an (maximum) operating pressure of2.1 MPa (300 lbf/in2) is given to meet the combined energeticrequirements.
Based on the requisite displaced volume for each actuator,Table VI summarizes the dimensions of each hand actuator (allof which are the same size), and shows the energetic output rel-ative to the requirements shown in Table V, where excess energy
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Fig. 16. Solid model of hand cylinders mounted to the distal forearm.
Fig. 17. Solid model of the full arm prosthesis.
(as given by a ratio greater than 1) will presumably overcomefrictional losses (e.g., tendon and cylinder seal friction), andprovide for additional performance.
Fig. 16 shows a solid model of the hand cylinders (Bimbamodel 011.5-DV) mounted on the distal forearm. The handcylinders, together with the cylinders for wrist flexion/extensionand abduction/adduction, form a group of seven actuators posi-tioned radially around and mounted to the distal forearm. Theactuator forces are applied through a series of tendons (700 Ntest sprectra cable) and cable sheaths. Cable tension is measuredfor each actuator by load cells (Measurement Specialties modelELFM-T2E-025L), which are located at the end of each pistonrod between the tendon and actuator.
VI. PROSTHESIS PROTOTYPE
The full prosthesis design is shown in Fig. 17. As describedin Section IV, the elbow cylinder is controlled by a servovalvemounted directly to the cylinder. The hand and wrist cylindersare controlled via eight servovalves, positioned in two banksof four valves each, as shown in Fig. 17. Gas flow from thevalves to the cylinders is routed via a network of stainless steeltubes, as shown in Fig. 18, which also serve to structurallysuspend the valve banks. Fig. 17 also depicts the propellant
Fig. 18. Gas control and distribution system. (a) Without the servovalve mo-tors. (b) With servovalve motors.
Fig. 19. Fully fabricated prosthesis prototype. See accompanying video fordemonstration of closed-loop controlled system.
cartridge, which is situated just proximal to the elbow joint. Thepropellant cartridge holds 200 mL of hydrogen peroxide, which,at a concentration of 70%, will provide 55 kJ of usable work.Although not shown in the model of Fig. 17, the propellant isrouted from the cartridge across the elbow joint through a pairof flexible fuel lines, each of which feeds a catalyst pack on theforearm. One of the catalyst packs supplies the elbow actuatorand measures 8 mm in diameter and 25 mm in length, while theother catalyst pack supplies the valve banks and measures 8 mmin diameter and 32 mm in length. Each catalyst pack is packedwith iridium-coated alumina granules, which convert the liquidhydrogen peroxide to gas.
Fig. 19 shows the completed prosthesis prototype, which,as shown (without the propellant cartridge), weighs 1.55 kg
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Fig. 20. Fabricated gas control and distribution system. (a) Without the ser-vovalve motors. (b) With servovalve motors.
(3.4 lb). The fabricated gas control and distribution system forthe wrist and hand actuators is shown in Fig. 20. Note that thepropellant cartridge has not yet been fully implemented, but hasbeen designed and fabricated, and when full will add 0.45 kg(1.0 lb) to the prosthesis for a fully charged prosthesis weightof 2.0 kg (4.4 lb).
Though not yet tested with on-board (propellant-based)power, closed-loop control and functionality of the prosthesishave been tested and demonstrated with an off-board supplyof cold gas (nitrogen). All nine actuators in the prosthesis arecontrolled by a linear proportional and integral (PI) force con-trol loop. Note that, as previously discussed, in the absence ofcontact, force control is mapped by compliance to motion con-trol in the hand, and thus, an explicit motion control loop is notrequired. The four direct-drive degrees of freedom (elbow andthree wrist) include explicit motion control loops, consisting ofproportional-integral-derivative (PID) loops wrapped around theinner force control loops. Representative responses of the outer-most loops for the prosthesis prototype are shown in Figs. 21–23.Specifically, Fig. 21 shows the responses for a 20◦ step com-mand for flexion and extension of the elbow joint; Fig. 22 showsthe responses for a 40◦ step command for flexion and extensionof the wrist flexion degree of freedom; and Fig. 23 shows theresponse for a 60 N step in force for flexion and extension ofthe index finger actuator in the absence of contact (i.e., whenthe finger is allowed to move in free space). Note that for the lat-ter, in particular, the shape of the step response differs in flexionand extension, since the former is forced by the actuator, whilethe latter is forced by the compliance of the finger joints. Notethat the inner force control loops around each cylinder largelyremove the frictional effects from the piston seals. Note also thatbacklash does not exist in the finger tendons. Backlash presentin the position-controlled degrees of freedom does not appearto be significant (even in the wrist pronation/supination degreeof freedom), presumably because all joints are essentially of thedirect-drive type (i.e., backlash is not multiplied through a highratio transmission).
An accompanying video (available at http://ieeexplore.ieee.org) demonstrates the mechanical aspects of the cold-gas-actuated prosthesis, and demonstrates its ability to perform var-ious tasks representative of activities of daily living. Two of
Fig. 21. Flexion and extension step responses for the elbow joint.
Fig. 22. Flexion and extension step responses for the wrist flexion degree offreedom.
Fig. 23. Flexion and extension step responses in force for the index finger inan unconstrained state (i.e., the force is resulting in a gross motion of the finger).
these tasks, throwing a ball and pouring a glass of water, arerepresented as a sequence of video frames in Figs. 24 and 25, re-spectively. Note that, for purposes of functional demonstration,motion and force commands to the arm were generated in ateleoperative manner by an operator wearing an arm exoskele-ton (not shown). Ultimately, for application to amputees, theauthors’ intention is to leverage the ongoing work in neuralprostheses (e.g., [1]–[14]) to generate motion and force com-mands directly from the amputee’s nervous system.
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Fig. 24. Video sequence of gas-actuated prosthesis throwing a ball.
Fig. 25. Video sequence of gas-actuated prosthesis pouring water.
VII. CONCLUSION
Gas actuators can provide a power density similar to hu-man skeletal muscle, but require both a viable source of on-board power and the existence of scale-appropriate servovalvesfor closed-loop control. The authors have proposed a solutionto the former, based on the use of the monopropellant hydro-gen peroxide, and have developed and demonstrated small-scaleservovalves that provide for the latter. Based on the monopro-pellant concept and these small-scale valves, the authors havedesigned, fabricated, and demonstrated (with cold gas) a 21 de-grees of freedom anthropomorphic transhumeral prosthesis withnine independent actuators. Current work is focused on transi-tioning from cold gas operation from an off-board source to hotgas operation from an on-board propellant cartridge.
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Kevin B. Fite (S’98–M’03) received the B.E., M.S.,and Ph.D. degrees in mechanical engineering fromVanderbilt University, Nashville, TN, in 1997, 1999,and 2002, respectively.
From 2002 to 2007, he was with the Departmentof Mechanical Engineering, Vanderbilt University,Nashville, TN, as a Research Associate. He is cur-rently an Assistant Professor in the Department ofMechanical and Aeronautical Engineering, ClarksonUniversity, Potsdam, NY. His current research inter-ests include power autonomous robot actuation, de-
sign and control of electromechanical and fluid power systems, and the designand control upper- and lower-extremity prostheses.
Thomas J. Withrow (M’07) received the S.B. de-gree in engineering science with specialization inbiomedical engineering from Harvard University,Cambridge, MA, in 2000, the M.S.E. degree inbiomedical engineering, the second M.S.E. degreein mechanical engineering, and the Ph.D. degreein biomedical engineering from the University ofMichigan, Ann Arbor, in 2001, 2002, and 2005,respectively.
He is currently a Postdoctoral Research Asso-ciate at the Department of Mechanical Engineering,
Vanderbilt University, Nashville, TN. His current research interests include thedevelopment, design, and testing of upper and lower extremity prostheses, softtissue biomaterial testing, injury biomechanics especially of the anterior cruci-ate ligament (ACL), and the biomechanics of injury prevention.
Xiangrong Shen received the B.E. degree in me-chanical engineering and automation from Shang-hai Jiao Tong University, Shanghai, China, in 1998,the M.S. degree in mechanical engineering from theUniversity of Nebraska–Lincoln, Lincoln, in 2003,and the Ph.D. degree in mechanical engineering fromVanderbilt University, Nashville, TN, in 2006.
He is currently a Postdoctoral Research Asso-ciate at the Department of Mechanical Engineering,Vanderbilt University.
Keith W. Wait (S’07) received the B.S. degreein mechanical engineering from Rice University,Houston, TX, in 2004. He is currently working to-ward the Ph.D. degree in mechanical engineering atthe Department of Mechanical Engineering, Vander-bilt University, Nashville, TN.
His current research interests include biologicallyinspired design and control of hexapedal robots.
Jason E. Mitchell received the B.S. degree fromTennessee Technological University, Cookeville, in1999, and the M.S. degree from Vanderbilt Uni-versity, Nashville, TN, in 2002, both in mechanicalengineering.
He is currently an R&D Engineer at theDepartment of Mechanical Engineering, VanderbiltUniversity, Nashville, TN. His current research in-terests include design and fabrication of upper andlower extremity prostheses.
Michael Goldfarb (S’93–M’95) received the B.S.degree in mechanical engineering from the Uni-versity of Arizona, Tucson, in 1988, and the S.M.and Ph.D. degrees in mechanical engineering fromMassachusetts Institute of Technology, Cambridge,in 1992 and 1994, respectively.
Since 1994, he has been with the Departmentof Mechanical Engineering, Vanderbilt University,Nashville, TN, where he is currently a Professor.His current research interests include the design ofhigh-energy-density robotic actuators, the control of
fluid-powered actuators, and the design and control of advanced upper and lowerextremity prostheses.
