AMAE Int. J. on Manufacturing and Material Science, Vol. 02, No. 02, May 2012  Control-Integrated Design by Theoretical Si...
AMAE Int. J. on Manufacturing and Material Science, Vol. 02, No. 02, May 2012Thus, the problem of multiple solutions of th...
AMAE Int. J. on Manufacturing and Material Science, Vol. 02, No. 02, May 2012In terms of these angles and the height of Po...
AMAE Int. J. on Manufacturing and Material Science, Vol. 02, No. 02, May 2012Now, (4c), (4d), (5a) and (6c) yield the coor...
AMAE Int. J. on Manufacturing and Material Science, Vol. 02, No. 02, May 2012acceleration along Bix direction correspond t...
AMAE Int. J. on Manufacturing and Material Science, Vol. 02, No. 02, May 2012In order to carry out the design, each basic ...
AMAE Int. J. on Manufacturing and Material Science, Vol. 02, No. 02, May 2012admissible design around           and respec...
AMAE Int. J. on Manufacturing and Material Science, Vol. 02, No. 02, May 2012For minimizing the deviation between the dema...
AMAE Int. J. on Manufacturing and Material Science, Vol. 02, No. 02, May 2012                                             ...
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Control-Integrated Design by Theoretical Simulation for a Torque-Actuated 6-SBU Stewart Platform

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A design algorithm has been proposed for a Stewart platform with six legs, each having a ball-screw at the middle and powered by a torque motor at the bottom. When a motor shaft rotates, the leg extends or collapses and the axis could rotate about a spherical joint supporting the motor. Consequent actuation from all the legs through a universal joint at the top of each causes the platform to change its pose. The joints at each end lie on the intersection of a pitch circle and a semi-regular hexagon. An inverse model that neglects friction and leg inertia has been employed in a step-by-step simultaneous search to determine the platform height at the neutral and
the radius of the bottom pitch circle within the constraint of permissible joint angle and motor specifications. The proposed control for a basic pose demand involves a feedforward
estimation of motor torque variation, a proportional-derivative feedback and appropriate compensating demand for minimizing unwanted coupled motion. The forward modeling of the pose dynamics and its Simulink implementation have
established the control as satisfactory.

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Control-Integrated Design by Theoretical Simulation for a Torque-Actuated 6-SBU Stewart Platform

  1. 1. AMAE Int. J. on Manufacturing and Material Science, Vol. 02, No. 02, May 2012 Control-Integrated Design by Theoretical Simulation for a Torque-Actuated 6-SBU Stewart Platform Biswajit Halder*, Rana Saha#, and Dipankar Sanyal+ Department of Mechanical Engineering, Jadavpur University, Kolkata, India * biswajeeet@gmail.com, #rsaha@mech.jdvu.ac.in, +dsanyal@mech.jdvu.ac.inAbstract—A design algorithm has been proposed for a Stewart refers to the prismatic joint within each leg and the lastplatform with six legs, each having a ball-screw at the middle alphabet S implies spherical joint with the moving platform atand powered by a torque motor at the bottom. When a motor the other end. Joints at the bottom and at the top are usuallyshaft rotates, the leg extends or collapses and the axis could arranged along the vertices of two regular [3] or semi-regularrotate about a spherical joint supporting the motor. Consequent hexagons [2, 4, 5]. Merlet [3] considered 3/6 configuration, inactuation from all the legs through a universal joint at the topof each causes the platform to change its pose. The joints at which one of the ends of two legs terminated to a common bi-each end lie on the intersection of a pitch circle and a semi- spherical joint, each located at the vertex of an equilateralregular hexagon. An inverse model that neglects friction and triangle. Controlling the input to each leg is necessary, causingleg inertia has been employed in a step-by-step simultaneous its length to change, with the objective of carrying the topsearch to determine the platform height at the neutral and platform through desired position and orientation, togetherthe radius of the bottom pitch circle within the constraint of called the pose.permissible joint angle and motor specifications. The proposed Liu et al. [2] and Merlet [3] modeled both the forward andcontrol for a basic pose demand involves a feedforward inverse kinematics of a Stewart platform and proposedestimation of motor torque variation, a proportional-derivative simplified solution schemes for the forward kinematics. Whilefeedback and appropriate compensating demand forminimizing unwanted coupled motion. The forward modeling the inverse kinematics deals with estimating the neutral lengthof the pose dynamics and its Simulink implementation have and stroke of the legs from the specified range of desiredestablished the control as satisfactory. platform pose, the forward or direct kinematics are meant for control analysis for predicting the platform pose from theIndex Terms—feedforward-feedback, forward modeling, known length of the legs. An ingenious solution scheme isinverse modeling, parallel manipulator, Simulink necessary for the forward dynamics problem to guide the mechanism through any desired instantaneous solution I. INTRODUCTION among the possible multiple solutions for an intermediate Parallel manipulators are widely used as laboratory-scale pose [7].flight or ship-motion simulators for assessing stability and Fichter [4] neglected the effect of leg inertia to obtain acontrol performance of certain on-board systems subjected simple forward dynamics model based on Newton-Eulerianto complex inertial loading. Any such manipulator has approach for computing actuating forces on a Stewart platformintegrated control for imparting a range of desired motions to corresponding to input of actuation forces to the legs.a large payload within a small workspace. In order to take Employing Newton-Eulerian analysis, Dasgupta andcare of the imprecision related to a number of passive joints Mrithyunjaya [5, 6] arrived at both inverse and forwardin the system and multiplicity of the response to a definite models for the kinematics and dynamics. Their predictioncommand, the control is critical. A Stewart platform [1] is the showed the effectiveness of a PD control for the force inputsmost popular parallel manipulator with six degrees-of- to the legs corresponding to different pose dynamics, withfreedom. the control estimated from the difference between the required A Stewart platform involves six linearly extensible legs and the predicted leg lengths at each instant of time. Sensorswith active electric or hydraulic drive for each. In a like LVDT is necessary to measure the instantaneous lengthconventional Stewart platform, the bottom end of each leg is of each leg in order to implement the internal control ofconnected by a spherical or universal joint to a stationary correcting the leg length leaving the pose control to be takenframe at the bottom and the top end is connected by a care of in an outer loop. A similar PD control strategy isspherical joint to the moving platform that supports a payload implemented using feedback of actuator length [8].on top. Legs only with prismatic joints between its upper and Recently, Andreff and Martinet [9] used a classicallower parts have been analyzed in the past, though nowadays perspective camera as an additional contactless sensor forvarious laboratories are exploring the use of ball-screw joints. constructing the platform pose directly and developed aBoth 6-SPS and 6-UPS configurations have been extensively control-devoted projective kinematic model. In theiranalyzed [2-6],where 6 stands for the number of joints of same integrated approach of designing the vision-control robot intype, the first alphabet S or U corresponds to spherical or the form of Stewart platform, both the internal and externaluniversal joint with the fixed frame at one end of each leg, P control loops have been considered together.© 2012 AMAE 13DOI: 01.IJMMS.02.02.51
  2. 2. AMAE Int. J. on Manufacturing and Material Science, Vol. 02, No. 02, May 2012Thus, the problem of multiple solutions of the forward parts, a motor, a spherical joint and another cylindrical stub.dynamics was overcome through additional sensor Fig. 1 shows System 1 in relatively greater detail with themeasurements. Tahri et al. [10] employed a central omni- motor indicated as M1, while in System 5 Points b5, B5, T5 anddirectional camera suitable for high-speed task. t5 mark the extent of the bottom stub, the leg and the top stub Design of a Stewart platform driven by controlled torque respectively. While T5 represents the universal joint and B5input from a variable-speed DC motor to each of the six legs the spherical joint, B5z shows the leg axis. For the other fivewith ball-screw joints has been reported here. Such a motor leg systems, the respective centerlines 1 to 4 and 6 representdrive has been chosen in view of the commercial availability the axes and filled circles the joints.of brushless DC motors and proven control performance of While the stubs on the top terminate at the points showneach with simple PID feedback. The active rotation of each by unfilled dishes at the periphery of the disc, the bottomleg about the motor axis has been considered in the modeling stubs end at a semi-regular hexagonal frame withthat did not arise in the earlier analyses with prismatic legs. circumscribing circle of radius rb and the center at o. TheBoth the inverse and forward models of the chosen frame fixes the mechanism on the ground. The open dishes atarrangement shown in Fig. 1 have been developed and the disc periphery also lie at the vertices of another semi-implemented in Matlab-Simulink frame leading to the control- regular hexagon.integrated design of the mechanism as a mechatronic system. While the universal joints Ti lie on a circle of radius rT The main objective of the proposed control design is to with center at Q, the spherical joints Bi lie on a circle of radiusobtain a control structure with minimal cross-coupled motion r with center at O. In the stationary and moving coordinateover the demanded basic motion in surge, sway, heave, roll systems with origins shownat Points o and p respectivelyor pitch. A thorough design analysis for the selection of the and the axes denoted by (x,y,z) and (px,py,pz) respectively,appropriate actuators has been accomplished in association the angular locations for Points biand ti can be expressed aswith optimizing the platform dimensions. Also, a controlstrategy along with control gains has been carried out. Ofcourse posing of compensating demands for reducing theunwanted cross-coupled motion is a notable feature of thecontrol strategy involving proportional and derivativefeedbacks and feedforward. where a letter outside and within parentheses in the subscript respectively identify a point and the origin or the axis direction of the coordinate system for the associated variable and half of t 1qt 2 is written as qs. At the neutral pose described in Fig. 1, Line oOQqp is vertical and the axis of each leg is collinear to its end stubs. The running length of the ith leg system can be determined as where superscript t, b and a to a variable stands for the top, bottom and actuated part of a leg. Particularly, at the neutral pose, indicated by a 0 in the subscript, axes of these parts are collinear. In other words, the running length at the neutral is equal to the distance of point bi and ti. Therefore, expressing Figure 1. Schematic of a Stewart Platform at Neutral Pose the coordinate of a point by a vector x, it can be written that II. PLATFORM CONFIGURATION Figure 1 shows the schematic of a Stewart platform The angular locations for Points Bi and for Points Ti in thesupporting a cylindrical payload on top of a circular disc stationary and the moving coordinates can be expressed aswith center at Point q and radius rt. However, the center ofmass of the payload together with the disc lies at Point p.Below the disc, there are six actuation systems i = 1 to 6.From the top towards the bottom, each system has a cylindricalstub, a universal joint, a leg comprising of upper and lower© 2012 AMAE 14DOI: 01.IJMMS.02.02.51
  3. 3. AMAE Int. J. on Manufacturing and Material Science, Vol. 02, No. 02, May 2012In terms of these angles and the height of Points p and q Only for pure yawing motion from the neutral, Points p, q and Q would remain stationary. For this case and the case of pureabove Point o at the neutral pose written as z p 0( o ) and heaving motion, the directions of axes pz and z remainz q 0(o ) respectively, the coordinates of points bi, Bi, Ti and ti coincident and the circles with Points Ti and ti at the peripherycan be written as would remain horizontal. The top stubs joining the respective peripheral points would undergo pure vertical motion during heaving and pure angular rotation during yawing. This angular motion is obviously equal to that of the payload about z-axis. The spherical joint permits this rotation. In the set-up being investigated, such joints have been considered at the bottom of each leg that would cause each leg to rotate passively also by the same amount. This yields a 6-SBU joint configuration, where B refers to the ball-screw joint replacing P for the conventional prismatic joint [5]. This configurationand has been taken up for the subsequent kinematic and dynamic modeling. Each leg in Fig. 1 has been chosen as a ball-screw joint. III. INVERSE KINEMATIC MODELINGThe lower part of the joint could be rotated by the motor An objective of the inverse kinematic modeling is tocoupled to it and the top part could extend or retract depending express the lengths of the six legs in terms of the pose. Whenon the direction of motor rotation. Of course, the rates and the mechanism is subjected to a pure surge, sway or heave,directions of the motor rotations together decide what type both the stationary and the moving coordinate systemsof motion the payload would undertake. This motion initiated remain parallel to each other and the correspondingfrom the neutral could be a pure translation or a pure rotation. ˆ component of the unit vectors e o and e p in these systems ˆWhile, the translations along x, y and z axes, respectivelyreferred as surge (Su), sway (Sw) and heave (H), describe the remain identical throughout the motion. It is conventional toposition acquired by Point p on the payload as describe a general angular motion as an ordered combination of roll, pitch and yaw. For any such motion, the two coordinate systems do not remain parallel any more. For a payload pose, the unit vectorsany rotational motion about x, y and z axes, respectively calledroll (R), pitch (P) and yaw (Y), provides the orientation of thepayload as and The direction B5z along which the translation of the upper can be related through the rotation matrix R p, o forpart of Leg 5 would take place has also been indicated in Fig. transforming a vector from the moving to stationary1. Thus for any motorized leg i, Biz is the direction along coordinate system [11] aswhich the leg length changes due to active motor rotationcausing motion of Ti relative to Bi. Accepting the direction Biyas the axis of rotation of the leg centerline, six rotatingcoordinate systems (Bix,Biy,Biz) are defined with origins at the wherestationary points Bi. For any pose away from the neutral, theaxes of the stubs and the leg of each actuation system doesnot remain collinear. Of course, the bottom stubs and Points o and O are withstationary. However, the axis Biz of each leg could rotate aboutthe spherical joint at Bi, causing angular deviation betweenthe axes of the bottom stub and the leg. Consequent to allthese rotations along with the translations of the upper partsof the legs, Points Ti, ti, Q, q and p would move. The universaljoint at each of Ti permits the deviation between the axes ofeach leg and the associated top stub, as necessitated tosustain the desired payload motion.© 2012 AMAE 15DOI: 01.IJMMS.02.02.51
  4. 4. AMAE Int. J. on Manufacturing and Material Science, Vol. 02, No. 02, May 2012Now, (4c), (4d), (5a) and (6c) yield the coordinates of Ti and ti andasand The velocity of Point Ti can then be obtained by using (4c), (11a) and (11b) as Hence, from (4a), (4b), (7a) and (7b), the fixed lengths of andthe ith top and bottom stubs, the instantaneous length of theith leg and the unit vectors along these can be determined as The velocity and acceleration given by (12a) and (12b) can be used to determine the velocity and acceleration of Points Ti in the rotating coordinate system for each actuator as and where the rotational matrixand There is a physical constraint on the maximum permissibleangle that a class of joint allows between its members. that couples the unit vector for the ith rotating coordinateTherefore, for the most demanding poses of the payload, the system defined asangles should be estimated at the spherical and the universaljoints between the axes of a leg and its associated bottomand top stubs. These angles can be found out as with that in the stationary coordinate system by the relation ˆ k Bi in (14b) defined by (8f) and the other two unit vectors inand (14b) obtained as It is also imperative to estimate the motor capacity interms of velocity and the acceleration demands arising from orthe desired payload poses. Using notations as explained inthe context of (1) and (2a), the linear and angular velocities andand the linear and angular accelerations of the payload canbe expressed respectively as along with the scalars in (14a) defined as It is evident that (14d) and (14f) have been defined so as to capture the velocity of ith leg in the plane BizBix. Since thethat in combination with (6c) yields upper part across the ball-screw joint has a universal joint at the top and the lower part has a spherical joint at the bottom, the motion of a leg can be described as follows. The first and second rates of extension of the actuated leg correspond respectively to the component of the linear velocity and acceleration of the top joint along the screw axis. Of course, the other nonzero component of the linear velocity and© 2012 AMAE 16DOI: 01.IJMMS.02.02.51
  5. 5. AMAE Int. J. on Manufacturing and Material Science, Vol. 02, No. 02, May 2012acceleration along Bix direction correspond to the angular IV. INVERSE DYNAMIC MODELINGmotion of the screw axis. In fact, the upper part and the motor The inverse model provides a way to estimate theplatform do not have any relative angular motion. These actuation input of current to the coils of the torque motors ui necessary for achieving the desired payload motion. Incomponents acquire the same swing velocity ω( Biy ) and addition to the motor-torque induced axial actuating force, to ui be estimated, is on the upper part of the ith screwacceleration  ( Biy) about the direction Biy and about the joint acting along the screw axis, the upper part receives ascrew axis Biz, the components undergo passive motion with transverse force as well due to the weight, friction ui ui and inertia forces corresponding to the rotation of the screw-rotational velocity ω( Biz ) and acceleration  ( Biz) that are joint axis about Point Bi.equal to the respective components of the payload rotation For the payload together with the disc with combined massabout the screw axis. Of course the extension rates of the leg p pare provided by the rotation of the lower part relative to the m p and the centroidal moment of inertia I ( px) , I ( py) pupper part of the screw joint with pitch p b . Thus for the ith and I ( pz ) , the equations of dynamics in the coordinate systemleg, using superscripts mi, ui and li respectively for the motor with origin at Point p shown in Fig. 1 can now be written asand the upper and lower parts of the screw joint, it can bewritten that with andand By neglecting the transverse force component, one can solve the above equations simultaneously with known description of the right-hand sides, so as to make an estimate of thewhere the component in (15c) to (15f) can be expressed as e required forces FTi( Biz) from motors. V. CONFIGURATION DESIGN OF A STEWART PLATFORM Sections 2 to 4 provide the equations for the length, velocity, force and joint angles pertaining to the legs of the mechanism depicted in Fig. 1 for a specified range of the basic displacements of a payload. The specifications have Equation (15a) provides the estimates of instantaneous been taken as ±0.3 m surge, sway and heave, ±150 roll and ±100 pitch and yaw for a 100 kg payload along with thelengths lia of all the legs with motorized actuation. Desired maximum limits for the rates given as ±0.05 m/s, ±12.60/s, 0.05lengths of the legs at the extremes can be found out from m/s2 and 1000/s2. Of course, deciding a suitable diameter forthese estimates obtained corresponding to the entire range the disc that would support the payload of a given size is theof desired output motion of the platform. starting point of the design. The diameter of the disc, the lengths of the top and bottom stubs and have been assumed to be equal to 0.5mm, 0.1m and 0.04m and 100 respectively. The design objective has been posed as estimating the feasible combination of , the neutral height of the disc from the base and , the base radius.© 2012 AMAE 17DOI: 01.IJMMS.02.02.51
  6. 6. AMAE Int. J. on Manufacturing and Material Science, Vol. 02, No. 02, May 2012In order to carry out the design, each basic motions has been The admissibility check to determine the limiting curveapplied such that the payload displacement from one end to corresponding to each constraint has been carried out bythe another in the specified range takes place in the shortest varying only one of the values of or at a time inpossible time for the specified limits of the velocity and the 0.01m step.acceleration. Danaher Motion EC2-B23-10L-05B actuatorswith 0.45m and 0.6m strokes have been considered, since the Out of the total set of values of and , nth value thdesired heaving by 0.6m demands a change of actuated length of and m value of can be represented by andof about 0.6m. A maximum joint angle, = 200 has been respectively which yields three limiting values- as maximum actuated length,considered as a design constraint for both the spherical anduniversal joints. Since the bottom joint angle has never been as minimum actuated length andfound to be more than the top one, only the top limit has as maximum top joint angle.been shown in Figs. 2 and 3. Two more design constraints and are used The values are represented by the following expressions:for the minimum and maximum running lengths limitsrespectively, estimated by (2) for an actuator at the minimumand maximum strokes. The lower constraining lengths havebeen taken as 1.18m and 1.286m for the actuators with0.45m and 0.6m stroke respectively and the correspondingupper constraining length for each has been obtained byadding the respective stroke. where Designing for the case of sway may be taken up for a sample illustration. For a chosen value of equal to 1.03mFigure2. Variation of limiting height-radius pair of platform shown in Fig. 2,, variation from 0.6m to 1.2m in 0.01m step may by solid, dashed and dash-dot lines respectively at the limits of be considered. Points a, b and c corresponds to 1.19m, 1.06mmaximum stroke, minimum stroke and maximum joint angle along and 0.62m respectively. It is evident that for any value ofwith the admissible zone indicated in shade for achieving prescribed below a and above both b and c are admissible from 6 DOF motion by Danaher-motion EC2-B23-10L-05B-450. the viewpoints of maximum limiting stroke, minimum limiting stroke and maximum joint angle respectively. In fact, the figure clearly reveals that any design in the range of between Points a and b for equal to 1.03m would support all the displacement variations excepting heave. The region shown by hatched lines within the limiting curves corresponding to maximum and minimum strokes for heave and sway displacements satisfies all the design objectives and constraints. Some of the limiting joint angle lines do not appear in Figs. 2 and 3 for being confined to region bounded by lower platform height and smaller bottom- circle radius than shown in the figures. Any value of the variable less than those on the solid lines and greater than those on the other two line types indicates admissible valuesFigure3. Variation of limiting height-radius pair of platform shown by solid, dashed and dash-dot lines respectively at the limits of corresponding to the constraint that a particular line typemaximum stroke, minimum stroke and maximum joint angle along represents. The variable pair prior to any step change thatwith the admissible zone indicated in shade for achieving prescribed made an admissible design inadmissible has been shown as 6 DOF motion by Danaher-motion EC2-B23-10L-05B-600. a point on a limiting line. It is apparent from Fig. 2 that the actuator with 0.45m stroke has a very narrow region of© 2012 AMAE 18DOI: 01.IJMMS.02.02.51
  7. 7. AMAE Int. J. on Manufacturing and Material Science, Vol. 02, No. 02, May 2012admissible design around and respectively equal friction and inertia [12]. Starting from the known initialto 0.95m 1.25m, whereas Fig. 3 shows that the actuator with conditions at the neutral, the platform pose can be determined0.6m stroke has a much wider region around 1.4m and 1.1m by simple integration ofrespectively. Hence, the dynamic admissibility test depictedin Fig. 4 has been carried out only for the latter actuator. in which the linear and angular velocity of the payload can be obtained by rewriting (16a) and (16b) in terms of variables at times t and t-t as andFigure 4. Variation of limiting height-radius pair of platform shown by solid, dashed and dash-dot lines respectively at the limits of peak force, minimum heave and maximum heave along with the admissible zone indicated in shade for achieving prescribed 6 DOF motion by Danaher-motion EC2-B23-10L-05B-600. Figure 4 depicts the dynamically admissibility of thedesign corresponding to the 100% duty-cycle force limitconstraint of 830 N available in case of Danahaer-motionEC2-B23-10L-5B-600 actuator along with keeping 7% marginfor friction and another 20% for other factors. In this figure,the critical cases of the kinematic-constraint admissibility zonehave also been plotted. It is evident that rb and z p0(o)respectively equal to 1100 mm and 1400 mm provide anacceptable design solution. The corresponding minimum, whereneutral and the maximum actuated lengths are 1286 mm, 1647mm and 1886 mm respectively. It may be mentioned at thisstage that the effect of the transverse inertial force in eachleg has been neglected in (16a) and (16b) so as to obtain sixequations with six unknown actuating forces on the movingplatform corresponding to its demanded pose variation. Of andcourse, (10) to (15d) should also be invoked during thecomputation. Besides the neglected inertia, the additionaleffects of joint friction and the longitudinal inertia of the toppart of each leg arise during the platform motion. These callfor a control strategy and its performance analysis throughtheoretical simulation of the forward modeling. in which the symbols have been used as follows – (a) , VI. FORWARD MODELING WITH CONTROL and are the torque coefficients of the universal, The forward modeling provides the variation of the pose spherical and ball-screw joints, (b) , and a r eof the payload with time corresponding to known variations the masses of the motor and the upper and lower parts ofof forces imparted to it. While the force component FTi Biz  each leg, (c) , and are the lengths of the motor and the upper and lower parts of each leg and respectively andalong the axis of Leg i originates from the motor,the (d) is the moment of inertia of the motorized leg about itscomponent FTi Bix perpendicular to it arises due to the centerline.© 2012 AMAE 19DOI: 01.IJMMS.02.02.51
  8. 8. AMAE Int. J. on Manufacturing and Material Science, Vol. 02, No. 02, May 2012For minimizing the deviation between the demanded variation of linear displacement have been added. The objective is toof the length and velocity of each leg from those mitigate the steady-state errors that have been predicted toestimated as and by a feedforward strategy, a arise in Case 1 by putting compensating demands as 51.7mmfeedback correction for the force has been considered as sway and 7.1mm heave for 150 roll and -34.6mm surge and 3.0mm heave for 100 pitch. Figs. 6(a) and (d) show that the basic dynamics for Cases 1 and 2 to be identical, whereas the predicted cross-coupled displacements in Figs. 6 (b), (c), (e)where the proportional and derivative gains are represented and (f) corresponding to Case 2 can be seen to be negligibleby kP and kD respectively. in the steady-state and quite small even during the transients. Cases 3 to 5 of Figs. 6(a) and 6(c) reveal identical dynamics VII. CONTROL DESIGN AND PERFORMANCE ANALYSIS for the demanded basic poses, if either the roll or the pitch is In order to accomplish the design analysis, the formulation applied as 100/s ramp demand for 1.5s or 1s respectively.presented above has been implemented in Simulink Case 3 correspond to the dynamics corresponding to thisframework. The objective of the study is to ascertain the tracking of the ramp demand is placed without anyfeasibility of the proposed control for achieving different compensation, the variations of the cross-coupledsteady demands of basic displacements starting from the displacement errors can be seen to be given by Case 3 isneutral pose of the platform in each case represented in Figs. each of Figs. 6 (b), (c), (e) and (f).5 and 6. While the maximum force limit of Danahaer-motion Though the predicted transient variations between CasesEC2-B23-10L-5B-600 actuator for 100% duty cycle of 830N 1 and 3 are different, the steady-state errors are almost equal.has been set as a constraint in the simulation, the values for In Cases 2 and 4, identical compensating demands have beenthe variables and parameters considered in the study are employed for mitigating the cross-coupled linear displacement errors. In both the cases, negligible steady-C b , C s and C u equal to 0.001, 0.002 and 0.0001 N-m/s state errors have been predicted. However, the transient errorsrespectively, , , and equal to 1.37×10-6, for Cases 2 and 4 have been found as negligible and86.54, 86.54 and 44.48 kg-m2 respectively , , significant respectively.and equal to 1.0, 0.179, 0.8 and 0.2m For Case 5, the compensations have been applied in raterespectively and , , , , equal to 3.5, 4.63 2.5 forms as 34.5mm/s sway and 4.7mm/s heave for roll demandand 200kg respectively. For the numerical simulation, the of 100/s over a period of 1.5s and -34.5mm/s surge and 3.0mm/proportional and derivative gains have been set as 500N/m s heave for pitch demand of 100/s over a period of 1.0s. Figs.and 120N-s/m respectively. 6 (b), (c), (e) and (f) for this case show only about 1.5mm and For the control of pure surge, sway or heave displacement, 0.75mm maximum transient cross-coupled heaves during rolla constant feedforward target of the actuated lengths and a and pitch dynamics respectively. Thus, a ramp demand of anconstant estimate of the motor torques corresponding to the angular displacement should be associated withtarget steady pose have been added to the feedback. Fig. 5 compensation in the form of rates.shows the predicted control performance to be quitesatisfactory. No significant cross-coupled displacement hasbeen predicted corresponding to these linear displacementdemands. Figure 6 depicts the predicted dynamics of the platformwith different control strategies employed for attaining twodifferent angular displacement demands. Figs. 6 (a) to (c)pertain to a 150 demand of roll, Figs. 6 (d) to (f) correspond toa 100 demand of pitch and the strategies have been numberedin each figure as 1 to 5. Case 1 shows the predicted dynamics,when the desired displacement has been applied as a simplestep demand. Though the basic demand can be seen as achieved byabout 1s, it is associated with cross-coupled steady-statedisplacement errors in both horizontal and vertical directions.In cases of roll and pitch demands, the coupled horizontalerrors per degree of the angular displacement have been Figure 5. Payload dynamics in surge, sway and heave for respectivepredicted as –3.5mm in sway and 3.5mm in surge respectively. step demandsThe corresponding heave errors are –0.5mm/0 and –0.3mm/0respectively.In each of Figs. 6 (a) to (f), Case 2 depicts thepredicted dynamics when besides raising the basic demandas a simple step in roll or pitch, compensating step demands© 2012 AMAE 20DOI: 01.IJMMS.02.02.51
  9. 9. AMAE Int. J. on Manufacturing and Material Science, Vol. 02, No. 02, May 2012 ACKNOWLEDGMENT We sincerely acknowledge RCI Hyderabad, India for a collaborating work with them and CSIR, New Delhi, India for the scholarship support. REFERENCES [1] D. Stewart, “A platform with six degrees of freedom”, Proceedings of Institute of Mechanical Engineering 180 (1) (1965) pp. 371–386. [2] K. Liu, J. Fitzgerald, F.L. Lewis, “Kinematic analysis of a Stewart platform manipulator”, IEEE Transactions on Industrial Electronics, 40 (2) 1993) pp. 282–293. [3] J.-P. Merlet, “Direct kinematics of parallel manipulators”, IEEE Transactions on Robotics and Automation, 9 (6) (1993) pp. 842- 845. [4] E.F. Fichter, “A Stewart platform-based manipulator: general theory and practical construction”, International Journal of RoboticsFigure6. Transient variations of (a) roll with (b) coupled sway and (c) coupled heave for 15 0 final demand and (d) pitch with (e) Research, 5 (2) (1986) pp. 157–82. coupled surge and (f) coupled heave for 10 0 final demand for [5] B. Dasgupta, T.S. Mruthyunjaya, “Closed-form dynamicdifferent control strategies – 1: only step demand, 2: step demand equations of the general Stewart platform through the Newton– with step correction, 3: only 10 0/s rate demand, 4: 10 0/s rate Euler approach”, Mechanisms ane Machines Theory 33 (7) (1998)demand with step correction and 5: 10 0/s rate demand with linear pp. 993–1012. displacement correction as rates. [6] B. Dasgupta, T.S. Mruthyunjaya, “A Newton–Euler formulation for the inverse dynamics of the Stewart platform CONCLUSIONS manipulator”, Mechanisms and Machines Theory 33 (7) (1998) pp. 1135–1152. A control-integrated six degree-of-freedom Stewart [7] M. Raghavan, “The Stewart platform of general geometry hasplatform with torque-motor operated six ball-screw actuators 40 configurations”, Journal of Mechanical Design, 115 (1993)has been designed. Both the inverse and the forward pp.277-282.modeling involving active rotation in the ball-screw actuators [8] C. Yang, Q. Huang, H. Jiang, O. Ogbobe Peter, J. Han, “PDhave been accomplished. Simulink implementation of these control with gravity compensation for hydraulic 6-DOF parallelmodels has provided the sizing as well as control design of manipulator”, Mechanisms and Machines Theory 45 (7) (2010)the system. pp. 666–677. A two-dimensional simultaneous step-by-step search has [9]N. Andreff, P. Martinet, Unifying kinematic modeling, identification, and control of a Gough–Stewart parallel robot into abeen executed for a given payload size and range of its vision-based framework, IEEE Transactions on Robotics anddemanded pose and the rates, along with the constraints of Automation, 22 (6) (2006) pp.1077-1086.the permissible maximum joint angles and available actuation [10] O. Tahri, Y. Mezouar, N. Andreff, P. Martinet, Omnidirectionalcapabilities in terms of stroke, force, velocity and acceleration visual-servo of a Gough–Stewart platform, IEEE Transactions onlimits. Control structures have been suggested for achieving Robotics and Automation, 25 (1) (2009) pp.178-183.any displacement demand in surge, sway, heave, roll or pitch [11] M.D. Shuster, “Spacecraft attitude determination and control”,with minimal cross-coupled dynamics. A PD feedback Chapter 5 in V.L. Pisacane and R.C. Moore (eds.), Fundamentals ofstructure has been integrated with the model-based Space Systems, Oxford University Press, 1994, pp. 245–336.feedforward estimates for all the cases. For the angular [12] P. Hamon, M. Gautier, P. Garrec, A. Janot, “Dynamic Identification of Robot with a Load Dependent Joint Friction Model”, in Proc. IEEE Conference on Robotics, Automation and Mechatronics, 2010, pp. 129-135.© 2012 AMAE 21DOI: 01.IJMMS.02.02.51

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