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Robotics » Robotics and Automation » "Parallel Manipulators, towards New Applications", book edited by Huapeng Wu, ISBN 978-3-902613-40-0, Published: April 1, 2008 under CC BY-NC-SA 3.0 license. © The Author(s).

Chapter 12

Dynamics of Hexapods with Fixed-Length Legs

By Rosario Sinatra and Fengfeng Xi
DOI: 10.5772/5434

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Overview

New hexapod design
Figure 1. New hexapod design
Kinematic notation of the ith leg
Figure 2. Kinematic notation of the ith leg
Geometry of the base and the moving platform
Figure 3. Geometry of the base and the moving platform
Motion Profile of moving platform
Figure 4. Motion Profile of moving platform
Actuator displacements
Figure 5. Actuator displacements
Actuator velocities
Figure 6. Actuator velocities
Actuator accelerations
Figure 7. Actuator accelerations
The computed actuator forces
Figure 8. The computed actuator forces
Kinematic mode
Figure 9. Kinematic mode
Model with counterweights mass
Figure 10. Model with counterweights mass
Balanced Hexapod using pantograph
Figure 11. Balanced Hexapod using pantograph
Leg counterweight
Figure 12. Leg counterweight
Mobile center of mass Hexapod
Figure 13. Mobile center of mass Hexapod
Fixed center of mass of Balanced Hexapod
Figure 14. Fixed center of mass of Balanced Hexapod
Graph for optimum design
Figure 15. Graph for optimum design

Dynamics of Hexapods with Fixed-Length Legs

Rosario Sinatra1 and Fengfeng Xi2

1. Introduction

Hexapod is a new type of machine tool based on the parallel closed-chain kinematic structure. Compared to the conventional machine tool, parallel mechanism structure offers superior stiffness, lower mass and higher acceleration, resulting from the parallel structural arrangement of the motion systems. Moreover, hexapod has the potential to be highly modular and re-configurable, with other advantages including higher dexterity, simpler and fewer fixtures, and multi-mode manufacturing capabilities.

Initially, hexapod was developed based on the Stewart platform, i.e. the prismatic type of parallel mechanism with the variable leg length. Commercial hexapods, such as VARIAX from Giddings & Lewis, Tornado from Hexel Corp., and Geodetic from Geodetic Technology Ltd., are all based on this structure. One of the disadvantages for the variable leg length structure is that the leg stiffness varies as the leg moves in and out. To overcome this problem, recently the constant leg length hexapod has been envisioned, for instance, HexaM from Toyada (Susuki et al., 1997). Hexaglibe form the Swiss Federal Institute of Techonology (Honegger et al., 1997), and Linapod form University of Stuttgart (Pritschow & Wurst, 1997). Between these two types, the fixed-length leg is stiffer (Tlusty et al., 1999) and, here, becoming popular.

Dynamic modeling and analysis of the parallel mechanisms is an important part of hexapod design and control. Much work has been done in this area, resulting in a very rich literature (Fichter, 1986, Sugimoto, 1987, Do & Yang, 1988, Geng et al., 1992, Tsai, 2000; Hashimoto & Kimura, 1989, Fijany & Bejezy, 1991). However, the research work conducted so far on the inverse dynamics has been focused on the parallel mechanisms with extensible legs.

In this chapter, first, in the inverse dynamics of the new type six d.o.f. hexapods with fixed-length legs, shown in Fig. 1, is developed with consideration of the masses of the moving platform and the legs. (Xi & Sinatra, 2002) This system consists of a moving platform MP and six legs sliding along the guideways that are mounted on the support structure. Each leg is connected at one end to the guideway by a universal joint and at another end to the moving platform by a spherical joint. The natural orthogonal complement method (Angeles & Lee, 1988, Angeles & Lee, 1989) is applied, which provides an effective way of solving multi-body dynamics systems. This method has been applied to studying serial and parallel manipulators (Angeles & Ma, 1988, Zanganesh et al., 1997) automated vehicles (Saha & Angeles, 1991) and flexible mechanisms (Xi & Sinatra, 1997). In this development, the Newton-Euler formulation is used to model the dynamics of each individual body, including the moving platform and the legs. All individual dynamics equations are then assembled to form the global dynamics equations. Based on the complete kinematics model developed, an explicit expression is derived for the natural orthogonal complement which effectively eliminates the constraint forces in the global dynamics equations. This leads to the inverse dynamics equations of hexapods that can be used to compute required actuator forces for given motions.

media/image1.png

Figure 1.

New hexapod design

Finally, for completeness of the dynamic study of the parallel manipulator with the fixed-length legs, the static balancing is studied (Xi et al., 2005).

A great deal of work has been carried out and reported in the literature for the static balancing problem. For example, in the case of serial manipulator, Nathan (Nathan, 1985) and Hervé (Hervé, 1986) applied the counterweight for gravity compensations. Streit et al. (Streit & Gilmore, 1991), (Walsh et al., 1991) proposed an approach to static balanced rotary bodies and two degrees of freedom of the revolute links using springs. Streit and Shin presented a general approach for the static balancing of planar linkages using springs(Streit & Shin, 1980). Ulrich and Kumar presented a method of passive mechanical gravity compensation using appropriate pulley profiles (Ulrich & Kumar, 1991). Kazerooni and Kim presented a method for statically-balanced direct drive arm (Kazerooni & Kim, 1990).

For the parallel manipulator much work was done by Gosselin et al. Research reported in (Gosselin & Wang, 1998) was focused on the design of gravity-compensated of a six–degree-of-freedom parallel manipulator with revolute joints. Each leg with two links is connected by an actuated revolute joint to the base platform and by a spherical joints the moving platform. Two methods are used, one approach using the counterweight and the other using springs. In the former method, if the centre of mass of a mechanism can be made stationary, the static balancing is obtained in any direction of the Cartesian space. In the second approach, if the total energy is kept constant, the mechanism is statically balanced only in the direction of gravity vector. The static balancing conditions are derived for the three-degree-of-freedom spatial parallel manipulator (Wang & Gosselin, 1998) and in similar conditions are obtained for spatial four-degree-of-freedom parallel manipulator using two common methods, namely, counterweights and springs (Wang & Gosselin, 2000).

In this chapter, following the same approach presented by Gosselin, the static balancing of the six d.o.f. platform type parallel manipulator with the fixed-length legs shown is studied. The mechanism can be balanced using the counterweight with a smart design of pantograph. The mechanism can be balanced using the method, i.e., the counterweight with a smart design of pantograph. By this design a constant global center of mass for any configurations of the manipulator is obtained.

Finally, the leg masses become important for hexapods operating at high speeds, such as high-speed machining; then in the future research and development the effect of leg inertia on hexapod dynamics considering high-speed applications will be investigated.

2. Kinematic modeling

2.1. Notation

As shown in Figure 2, this hexapod system consists of a moving platform MP to which a tool is attached, and six legs sliding along the guideways that are mounted on the support structure including the base platform BP. Each leg is connected at one end to the guideway by a universal joint and at another end to the moving platform by a spherical joint.

media/image2.jpeg

Figure 2.

Kinematic notation of the ith leg

The coordinate systems used are a fixed coordinate system O-xyz is attached to the base and a local coordinate system Ot-x t y t z t attached to the moving platform. Vector bi, si, and li are directed from O to Bi, from Bi to Ui, and from Ui to Si respectively. Bi indicates the position of one end of the ith guideway attached to the base, Ui indicates the position of the ith universal joint, and S i indicates the position of the ith spherical joint. Six legs are numbered from 1 to 6.

Furthermore, a local coordinate frame Oi-xiyizi is defined for each leg, with its origin located at the center of the ith universal joint. Two unit vectors are used. Unit vector uil is along the leg length representing the direction of the ith leg, and unit vector uis is along the guideway representing the direction of the ith guideway. The orientation of the ith coordinate frame with respect to the base can therefore be defined by a 3×3 rotation matrix, for i = 1,…,6, as

Qi=[uiauil×uiauil]
(1)

where uia is expressed as

uia=uis×uiluis×uil
(2)

Note that vector uil is configuration-dependent and determined for the given location of the moving platform; vector uis is constant and defined by the geometry of the hexapod.

For the purpose of carrying out the inverse dynamics analysis of the hexapod, the following symbols are defined. As shown in Figure 2, Ci is the center of mass of the ith leg, Cp is the center of mass of the moving platform, ci c˙i and c¨i are the position, velocity and acceleration vectors, respectively, of Ci with respect to the fixed coordinate frame, ρ¯ is the vector pointing from Ot to Cp with respect to the local coordinate frame Ot-xtytzt.

2.2. Kinematics

Consider one branch of the leg-guideway system, as shown in Figure 2, the following loop equation for i = 1,…,6, holds,

h+Rp¯ibisili=0
(3)

where h and R are the vector and rotation matrix that define the position and orientation of the moving platform relative to the base, respectively, p¯i is the vector representing the position of the ith spherical joint on the moving platform in the local coordinates.

Since the leg always moves along the guideway, si can be expressed as

where si is a scalar representing the displacement of the ith actuator along the guideway. Likewise, leg vector li can be expressed as

where l i is a scalar representing the fixed length of the ith leg. As mentioned in Section 2.1, the leg axis is parallel to the zi axis of the local coordinate frame Oi-xiyizi. In the light of eq.(1), uil can be expressed as

Substituting eqs.(4 & 5) into eq.(3) and rearranging it yields the following kinematics equations for the fixed-length leg hexapod, for i = 1,…,6,

siuis=h+Rp¯ibiliuil
(7)

To obtain the velocity of the moving platform, taking the time derivative of eq. (7) yields

s˙iuis=v+(ω×Rp¯i)(ωi×li)
(8)

where v and ω are the vectors representing the velocity and angular velocity of the moving platform, respectively, and ωi is the vector representing the angular velocity of the ith leg.

Furthermore, by taking dot product on both sides of eq.(8) by li, it leads to

s˙iuisli=[v+(ω×Rp¯i)]li
(9)

It is well known that the kinematic analysis of parallel manipulator leads to two Jacobian matrices, namely, the forward and the inverse Jacobian (Gosselin & Angeles, 1990). To find the Jacobians for the hexapod under study, rearranging eq.(9) yields the following form

s˙i(uisli)=[liT(Rp¯i×li)T]tp
(10)

where tp=[vT,ωT]T is the 6×1 twist vector of the moving platform. Consider all six legs it leads to the following expression

where s˙=[s˙1,...,s˙6] T is the 6×1 vector of the actuator speeds, and A and B are the 6×6 matrices representing the inverse and forward Jacobian of the hexapod and they are defined as

A=[l1T(Rp¯1×l1)Tl6T(Rp¯6×l6)T]
(12)
Β=diag(u1sl1,...,u6sl6)
(13)

Eq.(11) defines the differential relationship between the actuator speeds s˙ and the twist of the moving platform tp. Rewriting eq.(11) gives

Provided that B is invertible, the Jacobian matrix of the moving platform Jp can be given as

Jp=B1A=[Jp1T,...,Jp6T]T
(15)

where

Jpi=[liTliTuis(Rp¯i×li)TliTuis]
(16)

for i = 1,..,6. From eq.(14), tp can be expressed in terms of s˙ as,

where

To obtain the acceleration of the moving platform, taking the time derivative of eq. (14) yields

s¨=J˙ptp+Jpt˙p
(19)

where s¨=[s¨1,...,s¨6] T is the 6×1 vector of the actuator accelerations, t˙p=[aT ω˙T]T is the time derivative of the twist of the moving platform, J˙p is the time derivative of the Jacobian matrix of the moving platform obtained by differentiating Jp with respect to time, that is

J˙p=B1(A˙B˙B1A)
(20)

where A˙ and B˙ given as

A˙=[(ω1×l1)T((ω×Rp¯1)×l1+Rp¯1×(ω1×l1))T(ω6×l6)T((ω×Rp¯6)×l6+Rp¯6×(ω6×l6))T]
(21)
B ˙ =diag(u1s(ω1×l1),...,u6s(ω6×l6))
(22)

If the mass of the leg is uniformly distributed, then the center of mass is in its middle. The velocity of the center of mass can be given as

c˙i=s˙i+ωi×li2
(23)

Upon differentiating eq.(22), the acceleration of the center of mass can be given as

c¨i=s¨i+ω˙i×li2+ωi×(ωi×li2)
(24)

To obtain the leg angular velocity and acceleration, denote by Ei the 3×3 cross-product matrix associated with vector uil , then eq.(9) may be re-written as

Eiωi=1li[v+ω×Rp¯is˙iuis]
(25)

Consider all six legs, it forms a set of linear equations containing the unknowns of the leg angular velocity. There are three components of ωi for each leg. Because matrix Ei is a skew symmetric and singular, it is impossible to directly solve eq.(24). However, since the leg does not spin about its longitudinal axis, this indicates (Tsai, 2000)

In the light of eq.(25), eq. (24) may be rewritten as

where Ai is a 4×3 matrix and ei is a 4-dimensional vector, and they are defined as

ei=1li[v+ω×Rp¯is˙iuis0]
(29)

Solving eq. (26) leads to the expression for the leg angular velocity

ωi=lili2×[v+ω×Rp¯is˙iuis]
(30)

Now eq.(29) is substituted back into eq.(22), and the velocity becomes

c˙i=12[v+ω×Rp¯is˙iuis]
(31)

By examining eqs.(29) and (30), it may be noted that the two terms in the brackets are identical. The first term may be expressed as

v+ω×Rp¯i=[1,Epi]tp
(32)

where Epi is the cross-product matrix of Rp¯i . In the light of eq.(17), eq.(31) may be related to s˙ as

[1,Epi]tp=T1is˙
(33)

where 1 is the 3×3 identity matrix and T1i is the 3×6 matrix pertaining to the first term defined as

T1i=[1,Epi]Tp
(34)

The second term in eqs.(29) and (30) can also be expressed in terms of s˙

s˙iuis=T2is˙
(35)

where T2i is the 3×6 matrix pertaining to the second term defined as

T2i=[03,...,uis,...,03]
(36)

In eq.(35), 03 is the 3-dimensional null vector. The twist of the ith leg can be expressed in terms of s˙ as

where ti is the twist of the ith leg, i.e. ti=[c˙iT,ωiT]T , and the 6×6 matrix Ti is given as

Ti=[T1iTT2iT]T
(38)

Furthermore, the leg angular acceleration can be obtained by differentiating eq.(26) with respect to time, that is

Aiω˙i= e ˙ iA˙iωi
(39)

where

A˙iωi=[(ωi×uil)×ωi0]
(40)
e˙i=1li[a+ω˙×Rp¯i+ω×(ω×Rp¯i)s¨iuis0]
(41)

From eq.(38), vector ω˙i representing the angular acceleration of the ith leg is given as

ω˙i=1li2[(ωi×li)×(v+ω×Rp¯is˙iuis)+li×(a+ω˙×Rp¯i+ω×(ω×Rp¯i)s¨iuis)]
(42)

3. Dynamic modeling

3.1. The natural orthogonal complement method

Prior to performing dynamic modeling of the hexapod, a brief review of the natural orthogonal complement method (Angeles & Lee, 1988) is provided. Consider a system composed of p rigid bodies under holonomic constraints, the Newton-Euler equations for each individual body can be written, for i = 1,.., p, as

Mit˙i=WiMiti+wi
(43)

where ti is the twist of the ith body, wi=[niT,fiT]T represent the wrench acting on the ith body, ni and fi are the resultant moment and the resultant force acting at the center of mass. In general wi can be decomposed into working wrench wiw and non-working wrench wiN The former can further be decomposed as

wiw=wia+wig+wid
(44)

where wia,wig and wid are the actuator, gravity and dissipate wrenches, respectively.

In eq. (42), the 6×6 angular velocity matrix Wi and the 6×6 inertia matrix Mi are defined as

Wi=[ΩiOOO]Mi=[IiOOmi1]
(45)
, with
Ωi=(ωi×e)e
(46)

where Ii is the 3×3 matrix of the moment of inertia of the ith body, mi is the body mass, O denotes the 3×3 null matrix, and e is an arbitrary vector.

If consider all p bodies, the assembled system dynamics equations are given as

Mt˙=WMt+wW+wN
(47)

where the 6p×6p generalized mass matrix M and generalized angular velocity matrix W are defined as

Mdiag(M1,,Mp)
(48)
Wdiag(W1Wp)
(49)

and the 6p-dimensional generalized twist t, generalized working wrench wW and generalized non-working wrench wN are defined as

t[t1tp] wW[w1WwpW]
wN[w1NwpN]
(50)

It can be shown that the kinematic constraints hold the following relation with the generalized twist

where 06p is the 6p-dimensional null vector, K is the 6p×6p velocity constraint matrix with a rank of m which is equal to the number of independent holonomic constraints. The number of degrees of freedom of the system, i.e. independent variables, is determined as n = 6p - m. Denote the independent variables by s, they can be related to the twist as

t˙=Ts¨+T˙s˙
(53)

where T is a 6p×n twist-mapping matrix.

By substituting eq.(51) into eq.(50), the following relation can be obtained

where T is the natural orthogonal complement of K. As shown in (Angeles & Lee, 1988, 1989) the non-working vector wN lies in the null space of the transpose of T. Thus, if both sides of eq. (46) are multiplied by TT, in the aid of eqs. (51 & 52), the system dynamics equations can be obtained as

Is¨+Cs˙=TT(wa+wg+wd)
(55)

where the n×n generalized inertia matrix I and coupling matrix C are defined as

ITTMT ,
CTT(MT˙+WMT)
(56)

Furthermore, by defining the following generalized forces

τa=TTwa,τg=TTwgτd=TTwd
τI=Is¨+Cs˙
(57)

the inverse dynamics of the system can be given as

τa=τIτgτd
(58)

where τa is the vector representing the applied actuator forces.

3.2. Inverse dynamics

The key in applying the natural orthogonal complement method is to derive the expression for the twist-mapping matrix T, which relates the speeds of the independent variables to the generalized twist. For the hexapod under study, the independent variable s is the vector representing the actuator displacement, with the total number of six, as defined before. The generalized twist is expressed as

t=[t1t6tpc]
(59)

Note that t1 to t6 are the twists for the six legs. Since the twist in eq.(36) is defined at the center of mass of the leg, Ti represents the twist-mapping for the legs. For the moving platform, tpc is defined as the center of mass which may be expressed as

cp=h+Rρ¯
(60)

Differentiating eq.(59) gives

c˙p=v+ω×Rρ¯
(61)

In the light of eq.(60), the following relation can be obtained

tpc = Hp tp

where Hp is the 6×6 matrix defined as

In eq.(62), E is the cross-product matrix of Rρ¯ . Note that when ρ¯ is zero, i.e. the center of mass coincides with the coordinate origin, Hp becomes an identity matrix, and tpc = tp.

The twist-mapping matrix T for the hexapod under study can be given in the light of eqs.(17), (36) and (62) as

T=[T1T6HpTp]
(63)

where T is a 42×6 matrix. With T, the generalized forces can be defined according to eq.(56), and the applied actuator forces can be determined according to eq.(57).

4. Simulation

4.1. Geometric and inertial parameters

The geometry of the base and the moving platform is shown in Figure 3.

media/image142.png

Figure 3.

Geometry of the base and the moving platform

Accordingly, the coordinates of vector bi with respect to the fixed frame are given as

b1[Lb/2yb0]Tb2[(Lb+lb)/2,lbcs,0]Tb3[lb/2,(Lb+lb)csyb,0]Tb4[-lb/2(Lb+lb)csyb0]Tb5[(Lb+lb)/2,lbcsyb,0]Tb6[Lb/2,yb,0]T
(64)

where Lb and lb are the long and short side of the base hexagon, cs=cos(30) , and yb=(Lb/2+lb)tg(30) . Likewise, the coordinates of vector p¯i with respect to the local frame are given as

p¯1[Lp/2,yp0]Tp¯2[(Lp+lp)/2lpcs0]Tp¯3[lp/2,(Lp+lp)csyp0]Tp¯4[-lp/2(Lp+lp)csyp0]Tp¯5[(Lp+lp)/2,lpcsyp,0]Tp¯6[Lp/2,yp,0]T
(65)

where Lp and lp are the long and short side of the moving platform hexagon, and yp=(Lp/2+lp)tg(30) . The geometric parameters and inertial parameters are given in Tables 1 and 2, respectively. In Table 1, S is the guideway length, is the guideway angle between the guideway and the vertical direction, and l is the length of the leg. These three parameters are the same for all the guideways and legs. Parameters Lb, lb, Lp and lp are defined in Figure 3. In Table 2, m is the mass, and Ixx, Iyy and Izz are the moments of inertia.

S Guideway lengthY Guideway anglel Leg lengthL b Long side BPl b Short side BPL p Long side MPl p Short side MP
0.60 m45 o 0.50 m 1.00 m 0.09 m 0.50 m 0.09m 0.50 m1.00 m0.09 m0.50 m0.09m

Table 1.

Geometric parameters

m (kg)Ixx = Iyy (kg m2)Izz (kg m2)
Platform3.9830.0680.136
Leg0.3980.0474-

Table 2.

Inertial parameters

4.2. Numerical example

A simulation program has been developed using Matlab based on the method described in the previous sections. In terms of computation, as can be seen from eq. (54), the inverse dynamics of the hexapods mainly involves the twist-mapping matrix T and its derivative, which could be computed numerically for each time interval. This way, it is computationally more efficient. To further speed up computation, parallel computation techniques could be used. As shown in Figure 4, the motion part including actuator speeds and accelerations could be computed in parallel to the inertia part including mass matrix I and coupling matrix C. Since the program is done in Matlab, parallel computation is not realized. However, this strategy can certinaly be applied to model-based control using the dynamic equations.

In terms of singularity, as can be shown from eq. (63), the twist-mapping matrix T becomes degenerate when the moving platform Jacobian Jp (Tp = Jp -1) is singular.

The movement of the moving platform is defined in terms of 3-4-5 polynomials that guarantee zero velocities and zero accelerations at the beginning and at the end. The selection of a smooth motion profile is very important for the hexapod as it is operated under high speeds. The conventional machine tools are run at a maximum velocity of 30m/min with a maximum acceleration of 0.3 g. Hexapods can run at a maximum velocity of 100 m/min with a maximum acceleration over 1 g.

The first simulation is for high speed, with a maximum velocity of 102 m/min. The second simulation is for low high speed with a maximum velocity of 30 m/min. In both cases, the hexapod moves a distance of 0.1m along the z axis. The initial position of the moving platform is at xo = 0, yo = 0 and zo = 0.7m. Figure 4 shows the velocity profiles of the moving platform. Figures 5, 6 and 7 show the displacements, velocities and accelerations of the six actuators, respectively. Figure 8 shows the computed actuator forces. The simulations show that high speed motions result in large actuator forces.

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Figure 4.

Motion Profile of moving platform

media/image150.png

Figure 5.

Actuator displacements

media/image151.png

Figure 6.

Actuator velocities

media/image152.png

Figure 7.

Actuator accelerations

media/image153.png

Figure 8.

The computed actuator forces

5. Static balancing of the hexapod

The static and dynamic balancing is a classic problem in the theory of machines and mechanisms. In particular, when a mechanism is not statically balanced, the weight of linkage produces force or torque at actuators under static conditions and actuators have to contribute to support the weight of the moving links for any configurations. The problem becomes more serious for the parallel manipulator applied as flight simulator where the weight of the moving platform is very large with respect to the masses of the links. Static balancing also called gravity compensation is important. If the forces/toques exerted by joint actuators are reduced, the full potential of machine will be improved.

In this paragraph, following the same approach presented by Gosselin, the static balancing of the hexapod with the fixed-length legs is studied.

5.1. Static balancing using counterweight

The static balancing of the parallel manipulator under study is investigated using counterweights. The base coordinate frame Oxyz frame, is fixed to the base with Z-axis pointing vertically upward and the moving coordinate frame O'x'y'z' is attached to the moving platform. The Cartesian coordinates used to describe the pose of the platform are as shown in Fig. 9 given by the position of O' with respect to the fixed frame and the orientation of the platform represented by the rotation matrix Q

Q=[q11q12q13q21q22q23q31q32q33]
(66)

Using the counterweights, static balancing is obtained if the global center mass of the mechanism is kept stationary at any values of the independent variables. To choose an suitable constant, namely

media/image156.png

Figure 9.

Kinematic mode

where r is the position vector of the global mass center, and M is:

M=mp+i=16mi
(68)

where mp is the mass of the platform, mi is the mass of the leg. The global centre of the mass of the manipulator is written as

Mr=mprp+i=16miri
(69)

where rp is the platform center of the mass, ri is the leg center of the mass. From Fig. 9, vectors rp , ri can be derived, and substituted into eq.(69), yielding

Mr=mp(h+Qg)+i=16mi[(h+Qpi)(h+Qpibisi)lgili]
(70)

where g is the vector center of mass of the moving platform with respect to the frame O'x'y'z', h is the position of O' with respect to the fixed frame, pi is the position of the spherical joint with respect to the moving coordinate frame, bi is the position of the lower end of the guideway with respect to the fixed frame, li is the length of the leg, si can be written, for i=1,…,6, as

where s^i is the unit vector of guideway, ρi is the independent variable of the prismatic joint. In concise form, eq. (70) is expressed as

Mr=A1h+QB+i=16A5isi+A0
(72)

where

A1=mp+i=16mi(1lgili)
(73)
B=mpg+i=16mipi(1lgili)
(74)
A5i=milgili,fori=1,..,6 
(75)
A0=i=16A5ibi
(76)

The conditions for static balancing can be given for i=1,..,6, as follows:

A1=0,B=0,A5i=0,A0=0
(77)

From conditions A5i=0 i=1,..,6, one can obtain

By condition A1=0 , one can obtain

mp=i=16mi
(79)

Eq.(79) shows that the balancing by counterweight is impossible. If it was substituted in the condition B= 0,

mpg+i=16mipi=0
(80)

then one can obtain

g=i=16mipii=16mi
(81)

From eq. (81), it shows that the manipulator could be balanced by a device that provide a force that is

equal to the weight of the links and the platform;

in opposite direction of the weight;

5.2. Static balancing wit a pantograph counterweight

Since it is shown that the static balancing of the examined mechanism is impossible with the help of counterweights, we propose a method to add a pantograph connecting the moving platform O' to the fixed platform O, as shown in Fig. 10. The pantograph is a device that allows to keep two end points on the same line and keep their distance at the centre with a constant ratio. In this application it is possible to use a pantographs with two or more mesh as shown in Fig. 10 and Fig. 11, respectively. In both case the manipulator is balanced. The pantograph is fixed to the moving platform on the point O' by a spherical joint and fixed to the point O by an universal joint. The leg counterweight is shown in Fig.12.

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Figure 10.

Model with counterweights mass

media/image180.png

Figure 11.

Balanced Hexapod using pantograph

media/image181.png

Figure 12.

Leg counterweight

In this case, the mass M becomes,

M=mp+mp+ma+ma+i=16mi+i=16mi
(82)

where mp and mp are the mass of the platform and the mass of the platform counterweight, mi and mi are the mass of the legs and the mass of the legs counterweights, ma and ma are the mass of the pantograph and the mass of the of the pantograph counterweight. In this case, the global center of the mass of the manipulator is written as

Mr=mprp+mprp+mara+mara+i=16miri+i=16miri
(83)

where rp and rp are the platform center of the mass and the platform counterweight position, ri and ri are the legs center of the mass and the legs counterweight position, ra and ra are the pantograph center of the mass and the pantograph counterweight position. From Figs. 9-10, vectors rp , rp , ra ra , ri and ri can be derived and substituted into eq.(83), yielding

Mr=mp(h+Qg)+mp(h+Qg)+ma(hla|h|)+ma(hla|h|)++i=16mi[(h+Qpi)(h+Qpibisi)lgili]++i=16mi[(h+Qpi)(h+Qpibisi)lgili]
(84)

where, la is the center of mass of the pantograph with respect to the fixed frame, la * is the pantograph counterweight position with respect to the fixed frame, lgi is the length of the leg counterweight link, li is the length of the leg, si can be written, for i=1,...,6, as

In concise form, eq.(84) can be expressed as

Mr=A1h+QB+i=16A5isi+A0
(86)

where

A1=mp+mp+mala|h|+mala|h|+i=16mi(1lgili)+i=16mi(1lgili)
(87)
B=mpg+mpg+i=16mipi(1lgili)+i=16mipi(1lgili)
(88)
A5i=milgili+milgili,i=1,..,6
(89)
A0=i=16A5ibi
(90)

The conditions for static balancing can be given, for i =1,..,6, as follows

A1=0,B=0,A5i=0,A0=0
(91)

From conditions A5i=0 , for i=1,..,6, one can obtain

milgili+milgili=0
(92)

From eq. (92), for i=1,..,6, the following is obtained

lgi=milgimi
(93)

By condition A1=0 , i.e.,

mp+mp+i=16(mi+mi)+mala|h|+mala|h|=0
(94)

one can obtain

la*=|h|ma*(mp+mp*+i=16(mi+mi*)+mala|h|)
(95)

Finally, condition B= 0 leads to the following

mpg+mpg+i=16(mi+mi)pi=0
(96)

Eq.(96) shows that the static balancing can be achieved by fixing the global center of the mass of the moving platform, that of the legs and their counterweights at the same position, O'. In order to obtain it, the platform counterweight should be placed in the position:

g=mpg+i=16(mi+mi)pimp
(97)

Simulation is carried out to demonstrate the proposed method. The results are shown in Figs. 13-14, from which it can be seen that the centre of mass of the robot is non-stationary for non balanced case, while it is fixed for the balanced case.

After static balancing the global mass of the device increases by

ΔM=mp+ma+ma+i=16mi
(98)

The negative effect for the dynamic performance by the increasing global mass can be reduced by optimum design of the pantograph. A graph can be arranged to provide such help. Fig. 15 shows the ratios,

M+ΔMM,Ii+Ii*IiIa+Ia*Ia
(99)

which vary respect to the ratio ra */ h and lgi */ lgi and where Ii is the moment of inertia of the leg, Ii * is the moment of inertia of the leg counterweight whit respect of Pi, Ia is the moment of inertia of the moving platform and Ia * is the moment of inertia of the pantograph counterweight with respect of O. It should be noted that with a suitable design it is possible to reduce ΔM at the same time, it may increase Ii and Ia. The effect of gravity compensation on the dynamic performances was studied in detail in (Xi, 1999).

media/image221.png

Figure 13.

Mobile center of mass Hexapod

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Figure 14.

Fixed center of mass of Balanced Hexapod

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Figure 15.

Graph for optimum design

Input
Mobile platform
mass [kg]short side [mm]long side [mm]
8200800
Fixed platform 1
mass [kg]short side [mm]long side [mm]
/100400
Fixed platform 2
mass [kg]short side [mm]long side [mm]
/2501000
leg
mass [kg]l i [mm]l gi [mm]
0.5750375
Pantograph
mass [kg]side length [mm]r a [mm]
31000
Output
m a * [kg]m i * [kg]
171

Table 3.

Geometric and inertial parameters

7. Conclusion

In this chapter, the inverse dynamics of hexapods with fixed-length legs is analyzed using the natural orthogonal complement method, with considering the mass of the moving platform and those of the legs. A complete kinematics model is developed, which leads to an explicit expression for the twist-mapping matrix. Based on that, the inverse dynamics equations are derived that can be used to compute the required applied actuator forces for the given movement of the moving platform. The developed method has been implemented and demonstrated by simulation.

Successively, the static balancing of hexapods is addressed. The expression of the global center of mass is derived, based on which a set of static balancing equations has been obtained. It is shown that this type of parallel mechanism cannot be statically balanced by counterweights because prismatic joints do not have a fixed point to pivot as revolute joints. A new design is proposed to connect the centre of the moving platform to that of the fixed platform by a pantograph. The conditions for static balancing are derived. This mechanism is able to release the actuated joints from the weight of the moving legs for any configurations of the robot.

In the future research the leg inertia will be include for modeling the dynamics of the hexapod for high-speed applications.

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