AIAA Infotech@Aerospace Conference <br>and<br>AIAA Unmanned...Unlimited Conference
6 - 9 April 2009, Seattle, Washington
AIAA 2009-1870
Overview and Performance of the Front-End Robotics
Enabling Near-Term Demonstration (FREND) Robotic Arm
Thomas J. Debus1 and Sean P. Dougherty2
Alliance Spacesystems, LLC, Pasadena, CA 91103
On-orbit servicing of spacecraft offers the attractive possibility of extending the life of
existing military and commercial spacecraft. Several of the technical challenges and risks
associated with on-orbit servicing have already been mitigated during the Orbital Express
mission, in which a satellite outfitted with docking interfaces tailored to the specific mission
was successfully grappled. One of the key remaining technical challenges associated with
on-orbit servicing is the ability to service satellites that are not designed with custom
grappling interfaces, which constitute the majority of current on-orbit satellites. The goal of
the DARPA-sponsored FREND program is to overcome this challenge and develop,
demonstrate and ultimately fly robotic manipulator technologies designed to autonomously
perform unaided grappling. One of the key components of the FREND mission is the
development of a seven degree-of-freedom (DOF) flight robotic arm that can be used to
grapple, inspect, and service most of today’s spacecraft. This paper presents the capabilities
of such robotic arm and discusses its performance with respect to on-orbit servicing tasks.
I. Introduction
T
he promise of on-orbit robotic servicing has been heralded for nearly as long as NASA has existed as an agency,
as evidenced by illustrations by William Kocher in 19601. Actual on-orbit servicing of a spacecraft by a human
crew took place as early as 1973 with the repair of Skylab, and the benefits of such servicing have been
demonstrated more recently by the repair of the Hubble Space Telescope2. However, the true potential of robotic
servicing is only now beginning to be realized. The benefits of a robotic servicing spacecraft are manifold and
include such tasks as: repair, upgrade, rescue, refuel, repositioning, de-orbit, decommissioning, debris mitigation,
assembly of large structures (including antennae, telescopes, and solar arrays), sample return, and ferrying services
between different orbits or other orbital bodies such as the moon. An on-orbit servicing vehicle provides cost
savings to a customer by extending the life of a satellite as well as by providing a solution to the rendezvous and
docking problem; as such, the customer does not have to continually reinvent and re-fly these systems. These
capabilities enable potentially large market opportunities, and the commercial viability of on-orbit servicing has
been analyzed in many papers.2,3,4 The most immediate, economically-viable capability is likely to be the refueling
or repositioning of satellites that have run out of fuel; this capability is now practical using existing technologies.
The potential of on-orbit robotic servicing has, as of yet, been unrealized for two principal reasons: the
maturity/cost of the associated technologies, and the notion that satellites must be designed, from the very
beginning, to be serviceable. It is hard to make a practical case for on-orbit servicing if these two areas are not
addressed. However, once addressed, there is the potential for entirely new markets to emerge. To date, there have
been several missions to address the necessary technologies for robotic servicing including OMV, ROTEX, DART,
ETS-VII, and XSS-10/112. Most recently, DARPA’s Orbital Express (OE) program5 demonstrated the feasibility of
on-orbit servicing by successfully completing an end-to-end servicing mission including autonomous rendezvous
and docking, refueling, and component replacement. OE demonstrated that, for a prepared vehicle (i.e., a vehicle
designed to be serviced), there are now practical technologies available for reasonable costs. The final step is to
show that robotic servicing is possible even on existing or future satellites that were never intended to be serviced.
The DARPA-sponsored FREND program was created to prove the capability of autonomously executing an
unaided grapple of a spacecraft which was never designed to be serviced6. In successfully demonstrating this
capability in the Navy Research Laboratory’s Proximity Operations Test Facility, it addressed one of the few
1
2
Control Systems Engineer
Aerospace Engineer
1
American Institute of Aeronautics and Astronautics
Copyright © 2009 by Alliance Spacesystems, LLC. All rights reserved. Published by the American Institute of Aeronautics and Astronautics, Inc., with permission.
remaining obstacles to practical robotic servicing of spacecraft on orbit. Figure 1 shows the FREND system during
demonstrations in the Proximity Operations Test Facility. The FREND program developed and demonstrated a
flight robotic arm system with its associated avionics, end-effector, and algorithms. The robotic arm system,
developed by Alliance Spacesystems, was integrated into a high-fidelity testbed in which machine vision, trajectory
planning, and force feedback control algorithms were used to accomplish autonomous rendezvous and grapple of a
variety of spacecraft interfaces. Details of this robotic arm system and its associated capabilities are the focus of this
paper.
Photo Courtesy NRL
Photo Courtesy NRL
Figure 1. NRL On-Orbit Demonstration Testbed
The primary requirements of the robotic arm system are to provide high positional accuracy prior to the
grappling phase, compliance during the grappling phase, and the stiffness necessary to provide a rigid connection
between the arm and the serviced spacecraft. In addition, the arm must be highly dexterous to provide the
appropriate workspace needed for a wide range of tasks and satellite targets; must be able to support its own weight
during 1g testing; and must be capable of operating in thermal and radiation environments representative of geosynchronous orbits. The key requirements can be summarized as follows:
One side of the arm end-effector needs to fit inside a 3” radius, 40” deep cylinder
The arm needs to be designed to maximize dexterity
The arm must achieve a 17cm/s (6cm/s) velocity along its end-effector axis in 75% (95%) of its dexterous
workspace
The arm must weight less than 80Kg
The arm must support its own weight as well as a 5 Kg payload
The arm must survive launch loads
The arm is required to position a tool tip with a tip linear resolution of +/- 2mm and an angular accuracy of
+/- 0.4 deg
A central element of the FREND program is the robotic arm system developed to meet the above requirements.
This system includes a seven DOF flight robotic arm, a tool drive actuator and force/torque sensor at the end of the
arm, the low-level closed-loop algorithms, the avionics to control it, the thermal system, launch locks, power
system, spacecraft simulator, and other associated ground support equipment. The dominant design drivers are
stiffness, accuracy, packaging volume, and mass. The resulting system, shown in Fig. 2, is a two-meter class arm
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American Institute of Aeronautics and Astronautics
consisting of a three-axis shoulder joint and a three-axis spherical wrist connected midway by a revolute elbow. The
tool drive actuator provides an interface for actuating various end-effectors and the force/torque sensor is used to
sense contact and provide compliance once a grapple has taken place.
The resulting design offers many advantages in a fully flight-qualified package. It has a very high stiffness,
resulting in unmatched accuracy for a space robotic system. This accuracy can also be attributed to the highresolution/high-accuracy absolute position output sensors and near zero backlash at each joint. Comparable
systems, such as the robotic arms developed by Alliance Spacesystems for the Mars Exploration Rovers7 or Phoenix
Lander, have much lower stiffness and accuracy, as do the Space Station and Shuttle robotic arms. The FREND arm
is also 1-g testable in any orientation with a full 5Kg payload, another unique feature that greatly reduces test
complexity and allows for accelerated schedules. Because it was designed for space applications, it provides great
payload to mass capability compared to traditional industrial manipulators8. Its topology allows for a slender formfactor at the wrist, wide range of motion at each joint, outstanding dexterity, and the ability to fold back on itself and
stow in a small package for launch. Heritage9 actuator design also provides the strength, joint accuracy, and proven
reliability needed for the task. Finally, a unique cabling system developed by Alliance Spacesystems10 eliminates
the typical problems associated with round-wire harnesses such as snagging or added torques due to stiff harnessing.
This flex-harness, which services the entire arm, is similar to the flex-circuits technology used in laptop computers,
but on a much larger scale. Figure 2 shows the harness installed on the arm and also illustrates the cleanliness of
this cabling design.
Alliance Spacesystems developed two robotic arm systems for the FREND program: an EDU (Engineering
Development Unit) and an FPU (Flight Prototype Unit). Development was done on an accelerated schedule with the
fully integrated and tested EDU being delivered 13 months after program start and the FPU 9 months later. Because
the EDU arm showed exceptional performance (an accuracy of about 1mm at the tip, more than adequate torque and
speed margins, and a very capable/adaptable command and control scheme), the FPU arm varied only slightly from
the EDU. The primary difference on the FPU was the incorporation of real launch locks and the thermal control
system. The avionics remained the same for the EDU and FPU.
Figure 2. FREND Robotic Arm.
The remainder of this paper is arranged as follows. The next section provides a mechanical overview of the
system: in particular, packaging, dexterity, stiffness and mass, cabling and avionics are presented. Next, the
performance of the arm is discussed with regard to its accuracy and dynamic response. Conclusions are presented at
the end of the paper.
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American Institute of Aeronautics and Astronautics
II. Robotic Arm Design Overview
The proposed FREND arm is mechanically designed to optimize multiple criteria such as slenderness, mass and
stiffness. Some of these goals can be contradictory, and the design challenge is to find the right balance among all
the competing packaging, kinematics, structural and operation requirements. This section provides an overview of
these mechanical trade-offs.
A. Packaging
The packaging of the arm is driven by the end-effector (EE) slenderness requirement, the desire to maximize the
arm dexterous workspace, and the desire to minimize its stowed volume. To meet these requirements, the arm is
designed with an upper arm and forearm of equal lengths to maximize reachability. Offsets are added in the
shoulder, elbow and wrist pitch joints to maximize range of motion and minimize self-collision while providing
mechanical simplicity and easy cable management. The wrist area proved to be the most difficult to package due to
slenderness and dexterity requirements. The three wrist joints are arranged in a spherical configuration (i.e.,
intersecting rotation axes) to maximize dexterity; then, the wrist roll actuator is moved close to the elbow pitch joint
to ease cabling and satisfy the EE slenderness. Finally, an L-type bracket is used to connect the forearm to the wrist
pitch joint to minimize the distance between the wrist spherical center to the end-effector, therefore increasing the
stiffness and dexterity of the arm. Figure 3 illustrates the wrist packaging and shows how the wrist slenderness
requirement is met by allowing one side of the end-effector to fit inside a 3”(r) x 40”(h) cylinder.
Figure 3. Wrist Slenderness
Each joint’s range of motion is designed to be close to 360deg except for the wrist pitch and shoulder pitch. The
spherical wrist constraint (i.e., forearm collision) limits the range of motion of the wrist pitch to ±130°. The
Shoulder pitch range of motion is also limited to ±120° to limit possible collision with the arm mounting interface
(e.g., spacecraft deck, test stand). In addition, hard stops are added to each joint for safety and cable management
purposes and reduce the range of motion by +/-10°.
The design of the arm stowed configuration during launch is both a function of packaging volume requirements
and maintaining acceptable loading during the launch phase of the mission, potentially the most severe environment
the arm will experience. The optimum stowed configuration, shown in Fig. 4, allows the arm to be packaged inside
a 1.4m (L) x 1m (W) x 0.65m (H) box while minimizing the launch load. The foremost challenge when designing
launch locks is to protect the mechanisms and structure while not over-constraining the arm so as to preclude
thermal expansion/contraction mismatch driven damage or failed release. To do so, a semi-kinematic restraint
geometry is utilized with over-constrained points arranged so that lines of force are aligned along acceptably
compliant loading directions of the arm structure. The launch restraint system for the FREND arm is challenging
since there are seven DOF and thus seven separate actuators to protect from launch loads. The arm requires four
locks during launch. A six degree of restraint (DOR) lock is needed to secure the shoulder yaw, a two DOR
mechanism is needed to secure the elbow, a one DOR is required to lock the wrist pitch, and a four DOR is needed
for the tool drive. Nitinol-triggered pin-pullers are baselined as release mechanisms. The most challenging launch
lock is by far the one associated with the tool drive. In order to avoid overloading of the wrist geartrain and
bearings, the tool drive lock is designed to react the launch loads in the radial directions as well as the moments
around these directions. The tool drive is locked inside a U-shaped cradle using a pin puller. The cradle is allowed
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American Institute of Aeronautics and Astronautics
to rotate around its axis using four rollers mounted on the base of the launch lock and a set of cross-roller rails is
also used to permit linear motion of the base.
Figure 4. FREND Arm Stowed Configuration
B. Dexterity
To assess the arm’s dexterity, the existence of joint solutions corresponding to points sampled uniformly inside
the spherical Cartesian workspace volume is analyzed. Using the natural symmetry of the arm, the workspace
volume is reduced to that shown in Fig. 5. The dexterity of the arm is analyzed by extracting uniformly spaced
points across the workspace. The possible orientations associated with each end-effector point are then defined by a
uniformly sampled hemisphere around the EE point. A point is considered dexterous if at least 75% of its associated
orientations result in a feasible arm pose (i.e., a pose that satisfies the joint limits, precludes self-collision, and is not
singular). Due to the joint redundancy of the FREND arm, the augmented pseudo-inverse, J+, is used to compute
the Cartesian-to-joint mapping as shown in equation (1). The pseudo-inverse uses a damping factor α to avoid
discontinuity near a singularity and a matrix W to weight the joints according to a desired objective. The matrix W
is set to be the arm inertia matrix evaluated at the arm nominal pose to penalize high velocities associated with the
high inertia joints. To avoid non-cyclic behavior (i.e., a closed path in Cartesian space resulting in an open path in
the joint space), the pseudo-inverse also uses the joint null-space to optimize a secondary objective function11. The
FREND arm uses principally joint limits as the principal and collision avoidance as the secondary optimization
criterion. The integration of the rate controller expressed in (1) is used for position control.
(
)
q& = J + X& + I − J + J Z q&
(
J + = W −1 J T JW −1 J T + α 2 I
)
−1
(1)
The results of the dexterity analysis are illustrated in Fig. 5. The points successfully satisfying the dexterity
criteria defined above are green colored whereas the other points are color coded from yellow to red with decreasing
percentage of dexterity. The analysis shows that the arm is the most dexterous inside the workspace delimited by
the elbow and the wrist; these results are analogous to that of the human arm.
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American Institute of Aeronautics and Astronautics
100
90
80
70
60
50
40
30
20
10
0
Figure 5. Arm Dexterous Workspace
The augmented Jacobian is also used to flow down the Cartesian velocity requirements to joint velocity
requirements. Possible tip velocity orientations associated with each position/orientation point in the dexterous
workspace are represented using a uniformly discretized sphere. Equation (1) is then used to compute the joints
velocity corresponding to these Cartesian velocities. This analysis shows that joint velocities of 1.3 RPM for the
shoulder yaw and pitch, 2.5 RPM for the shoulder roll and elbow pitch, and 4 RPM for the wrist joints are needed to
meet the required tip linear velocity magnitudes of 17cm/s in 75% and 6cm/s in 95% of the dexterous workspace.
These joint velocities also result in tip angular velocity magnitudes of 0.6 rad/s in 75% and 0.2 rad/s in 95% of the
dexterous workspace.
C. Actuator Architecture
To achieve the desired +/-2mm linear positional accuracy requirement, gears and ball bearings, the “soft”
components of the arm are carefully selected to minimize mass while achieving high levels of stiffness. To do so,
all joint actuators except the tool drive are based on similar gearing architecture9 in which the motor torque is first
increased using planetary gears and then stepped-up using harmonic gearing. This method combines the high
efficiency of the planetary gears with the excellent stiffness-to-weight ratio (specific stiffness) of the harmonic gear.
Using harmonics as a final reduction pass substantially reduces the compliance and backlash effects associated with
planetary gearing. The mass of the FREND arm is 77kg including cabling and launch locks and excluding the 5kg
end-effector payload and the drive electronics. Accuracy and the need to test the arm in 1g drive the size of the
harmonics and the bearings, which in turn drives the size and mass of the housings and output shafts. The housings
and tubes fittings are made of titanium to match the coefficient of thermal expansion (CTE) of the gears and
bearings and meet the -55C survival temperature requirement. All of the actuators and tube fittings are also
equipped with a thermal control system (TCS) to operate in geostationary orbit, consisting of heaters, platinum
resistance thermometers (PRT) and thermostats. In addition, thermal blankets are also needed to provide increased
insulation.
The motors are selected to meet the velocities discussed in the previous section and the torques required to move
a 5Kg payload in a 1g environment. The low operational velocities and accelerations combined with the high gear
ratios minimize the effects of output inertia as well as the Coriolis and Centrifugal effects. As a result, the
operational torques are mostly driven by gravity and friction effects, especially at low temperatures. Friction-based
power off brakes are added to the motor to provide a safety mechanism during 1g testing and a way to efficiently
(i.e., minimal power) rigidize the arm after the grapple phase in 0g. The brake holding torques are selected such that
they can be overpowered by the motor if necessary.
D. Harnessing
All power and signal cables that traverse the arm are contained within a flex-cable which is manufactured using
printed circuit board methods as shown in Fig. 6. At each joint, the cabling is configured in a “clock spring” shape
to dynamically bend without imposing small bend radii or developing large bending strains. This shape allows for
large ranges of motion of the joints and is a function of the allotted annular space for the expanding and contracting
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American Institute of Aeronautics and Astronautics
clock spring. The size of the spools and the desired arm range of motion drive the length of the cables. A maximum
cable length of approximately 11m is required to connect the arm end-effector to the connector bulkhead located
near the shoulder yaw. In total, over 70 meters of flex-cable are needed for the entire arm and more than 500 traces
are used to carry the power and telemetry across the harness. The location, sizing and spacing of these traces is a
challenging task that requires several iterations to minimize noise and capacitive coupling. This effort is particularly
important at the end of the arm where video signals need to be carried over the maximum cable length. Multi traces
per phase are needed to carry the high current required to drive the actuators, and printed circuit boards are utilized
to simplify the routing between these power traces and the motor wires. Finally, Micro-D connectors are used to
connect the ends of the flex-cable to the arm bulkhead. The flex-cable allows this complex wiring to be managed
very effectively and packaged cleanly without the need for large service loops. In addition, this cabling
configuration allows for easier assembly and inspection, and, if necessary, disassembly and repair.
Figure 6. Flex Cabling Configuration of the FREND Arm (left) and Flex Material Sample (right)
E. Avionics and Software
In addition to the novel mechanical design developed for the FREND arm, custom electronics and software were
also developed to drive the arm and handle the kinematic redundancy associated with its 7 DOF. FPGA-based
motion control boards (MCB) are used for low level closed-loop position control of the joints. The MCBs are
custom designed to provide the flight-like avionics needed to drive the actuator brushless DC motors and read the
position sensors. These 3U form factor boards are enclosed in a 30cm (L) x 19cm (D) x 12cm (H) enclosure fitted
with a heat sink to provide convective cooling during 1g testing as shown in Fig. 7. Note that analysis has shown
conductive cooling to be adequate on orbit. Each board communicates to the electronic ground support equipment
(EGSE) through an RS422 interface. Commands and telemetry are sent and received at 500Hz for the most critical
data. In addition, low priority telemetry can also be received at 1Hz.
Figure 7. FREND Avionics
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American Institute of Aeronautics and Astronautics
A ground support equipment unit, shown in Fig. 8, consists of the arm power supplies, a test conductor ‘ground
station’, a Telemetry Gateway computer and an embedded rack-mounted ‘Spacecraft Simulator’ computer, which
are used to drive the arm. The VxWorks based Spacecraft Simulator is in charge of the real-time motion and torque
planning of the arm, inverse kinematics, fault monitoring and protection. All telemetry received by the Spacecraft
Simulator is passed to the Telemetry Gateway, where it is recorded by a RAID storage array. Here, telemetry
packets are filtered at a user-defined, real-time adjustable rate before being passed to the ground station. At the
ground station, the filtered telemetry is de-commutated and displayed via an SCL command line interface and/or
custom Labview windows. The SCL interface allows for scripting of regularly used command processes, while the
Labview interface provides a flexible means of data display enabling features such as real-time plot construction
from multiple telemetry points, mathematical manipulation of telemetry, signal analysis, etc. Commands can be sent
from either interface directly to the arm’s MCBs, or to the real-time control algorithms located on the Spacecraft
Simulator. The direct command interface to the MCBs allows for open-loop or closed-loop joint commanding as
well as MCB configuration commands (to set control loop parameters, fault masking, etc.).
The motion planner implemented inside the Spacecraft Simulator allows the arm to be operated either using joint
commands or tip commands. In addition, files containing pre-planned joint trajectories can also be sent directly to
the MCBs from the ground station. Torque computing is also implemented inside the simulator to compensate for
gravity, friction and varying inertia and therefore improve the dynamic response of the arm while operating in 1g.
Multiple fault flags can be set to ensure hardware and personal safety when working around the arm. In particular,
current limits, velocity limits, command limits and position limits (i.e., soft stop) are constantly monitored for
safety. In addition, two emergency modes are available to stop the arm: an ‘emergency hold’ switch is available to
stop the arm while maintaining its servo-controlled position, and an ‘emergency stop’ is available to stop the arm by
directly cutting its power.
Figure 8. FREND Electrical Ground Support Equipment
III. FREND Arm Performance
The FREND robotic arm has been tested at multiple levels to verify whether it met all the requirements
presented in section I. The arm was 1) tested at the part level to characterize the performance of its components
including motors, sensors and harness, 2) tested at the subsystem level to evaluate its actuators performance, and 3)
tested at the system level to verify its performance against the requirements. System level testing was primarily
conducted at ambient temperature and pressure with the exception of thermal-vacuum testing. These tests included
trajectory accuracy, settling time, calibration repeatability, vibration, thermo-vacuum and EMI/EMC. This section
presents the results associated with the arm settling time and accuracy tests.
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A. Sampling Time Test
The purpose of the settling time test is to characterize the natural frequency and damping ratio of the arm, to
validate the mass-stiffness model of the arm and understand its dynamic limitations. The test is performed with the
arm posed in the configuration that produces the longest settling time following the input perturbation: arm fully
extended with a 5Kg end-effector mass simulator attached to its wrist. The perturbation is produced by carefully
cutting a 10Kg mass suspended by a nylon line located directly below the wrist. Figure 9 shows the decay envelope
measured by a laser tracker. A nonlinear least squares fit assuming an underdamped second order system response
is used to determine the natural frequency and damping ratio of the system as shown in Fig. 9. The estimation
indicates that the arm possesses a natural frequency in its extended pose of 5.2Hz and a damping ratio of 1.7%
resulting in an estimated 2% settling time of 7.2 seconds.
4
x 10
-3
Measured vs. Simulated Responses
3
Experimental Data
Estimated Data
Amplitude X-dir (m)
2
1
0
-1
-2
-3
-4
2.5
3
3.5
4
4.5
Time (sec)
Figure 9. Impulse Response of the Arm
B. Relative Accuracy Test
The objective of the accuracy test is to characterize the tracking errors associated with representative on-orbit
arm trajectories. Relative tip accuracy of the arm is computed by mapping the joint errors to tip position and tip
orientation errors using the forward kinematic model of the arm. Relative accuracy is of greater concern than
absolute accuracy for the FREND arm since the majority of the uncompensated absolute errors (e.g., flexibility,
calibration, 1g sag) can be taken out using the camera-based tracking system.
Figure 10-13 illustrates the arm’s accuracy for various arm on-orbit moves before grappling. The trajectory is
divided into seven segments as shown in Fig. 10. The first segment (5s to 45s) corresponds to a self motion of the
arm in which the arm starts in ‘V’ configuration similar to the one shown in Fig. 2 and then rotates 90 degrees
around its shoulder-wrist axis, as shown in Fig.11, while maintaining its tip position and orientation fixed. This type
of motion, unique to arms with seven DOF or more, can be used for collision avoidance and also illustrates the
arm’s unique dexterity. The second segment of the trajectory (45s to 72s) corresponds to a 90 degree sweep of the
wrist’s yaw at 8deg/s. The third segment (72s to 86s) corresponds to a 6cm/s fast traverse of the arm that simulates
a fast plunge toward the target satellite. The fourth segment (86s to 112s) corresponds to the search phase in which
the arm scans the target satellite for grappling features (e.g., bolt hole, Marmam ring). The fifth segment (112s to
135s) corresponds to a 1cm/s slow forward plunge towards the detected target. The last two segments of the
trajectory represent two possible wrist motions. Segment 6 (135s to 151s) illustrates a wrist elevation move in
which the wrist pitch is rotated 30 degrees, and segment 7 (151s to 177s) illustrates a combined wrist yaw and wrist
roll motion corresponding to an azimuth motion of the wrist.
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0.06
Tip Velocity Magnitude(m/s)
0.05
0.04
0.03
0.02
0.01
2
1
0
0
20
40
60
3
4
80
100
Time (s)
5
120
6
7
140
160
180
100
Time (s)
Wrist Roll
200
100
0
-100
-200
0
100
Time (s)
200
-100
0
100
Time (s)
Wrist Pitch
200
0
-50
-100
0
100
Time (s)
200
Shoulder Roll
50
0
-50
-100
0
100
Time (s)
Wrist Yaw
200
0
100
Time (s)
200
Joint Position (deg)
0
-50
Joint Position (deg)
0
Shoulder Pitch
0
Joint Position (deg)
50
Joint Position (deg)
Shoulder Yaw
100
Joint Position (deg)
Joint Position (deg)
Joint Position (deg)
Figure 10. Seven Segment Arm Trajectory – Tip Velocity Magnitude
Elbow Pitch
150
100
50
0
100
Time (s)
200
0
-100
-200
Figure 11. Seven Segment Arm Trajectory – Joint Position
The commanded and measured tip positions corresponding to the seven segment trajectory are shown in Fig. 12.
The tip position is shown in the arm base frame in which the X axis is aligned with the gravity vector, the Z axis is
aligned with the shoulder yaw axis, and the Y axis completes the orthogonal frame. The tracking error results
presented in Fig. 13 show that the tip error is less than 1mm for most of the trajectory including transients between
the segments. Along track positioning errors are the largest whereas cross-track errors (e.g., YZ tip error during X
tip move) are the smallest (below half a millimeter). The largest error occurs along the track of the 6cm/s fast
plunge that occurs between the 72 and 86 seconds. This motion is mostly a pitch move as illustrated in Fig. 11 and
inspection of the joint errors shows that this error is mostly due to a 0.1deg following error at the elbow joint.
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X-Tip Position (m)
0.5
0
-0.5
0
20
40
60
80
100
120
140
160
180
Y-Tip Position (m)
Time (s)
1
0.5
Commanded Tip Position
Measured Tip Position
0
-0.5
0
20
40
60
80
100
120
140
160
180
100
120
140
160
180
Z-Tip Position (m)
Time (s)
2.5
2
1.5
1
0
20
40
60
80
Time (s)
Figure 12. Arm Commanded Vs. Measured Tip Position
-3
X-Tip Error (m)
1
x 10
0.5
0
-0.5
-1
0
20
40
60
80
100
Time (s)
120
140
160
180
20
40
60
80
100
Time (s)
120
140
160
180
20
40
60
80
100
Time (s)
120
140
160
180
Y-Tip Error (m)
-4
15
x 10
10
5
0
-5
0
-4
Z-Tip Error (m)
5
x 10
0
-5
-10
-15
0
Figure 13. ARM Relative Accuracy
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IV. Conclusion
The design of a space flight robotic arm tailored to on-orbit servicing missions presents a significant challenge
requiring that many parameters (e.g., stiffness, mass, packaging, dexterity, etc.) be successfully met simultaneously.
The FREND arm was built to meet these challenges by using an iterative design process in which these key
parameters were optimized to achieve and exceed the desired task requirements. This process resulted in a seven
degree-of-freedom, two-meter flight arm with high dexterity and high accuracy unique to the design. In addition, a
novel set of flight-like avionics developed for the FREND program offers the power and versatility required to
implement the complex control algorithms needed to perform increasingly demanding space robotic tasks.
The robotic arm has been successfully subjected to vibration, thermal and EMI/EMC tests (i.e., to simulate
launch and space environments) during the FREND program. It has been extensively used by NRL and has proven
to be very reliable and capable for performing on-orbit servicing tasks. The arm is a new benchmark for space-rated
robotic arms and provides the performance and flexibility required for future on-orbit servicing missions.
Acknowledgments
The authors would like to thank the Navy Research Laboratory and the Defense Advanced Research Project
Agency for their support.
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