Design of Minimally Actuated Legged Milli-Robots Using Compliant Mechanisms and Folding by Aaron Murdock Hoover a Dissertation S

Design of Minimally Actuated Legged Milli-Robots Using Compliant Mechanisms and Folding

Aaron Murdock Hoover

A dissertation submitted in partial satisfaction of the requirements for the degree of

Doctor of Philosophy in

Engineering - Mechanical Engineering

in the

GRADUATE DIVISION

of the

UNIVERSITY OF CALIFORNIA, BERKELEY

Committee in charge:

Professor Ronald S. Fearing, Co-chair Professor Dennis K. Lieu, Co-chair Professor Albert P. Pisano Professor Carlo H. Sequin´

Spring 2010 Design of Minimally Actuated Legged Milli-Robots Using Compliant Mechanisms and Folding

Aaron Murdock Hoover Abstract

Design of Minimally Actuated Legged Milli-Robots Using Compliant Mechanisms and Folding

Aaron Murdock Hoover Doctor of Philosophy in Engineering - Mechanical Engineering

University of California, Berkeley Professor Ronald S. Fearing, Co-Chair Professor Dennis K. Lieu, Co-Chair

This thesis explores milli- and meso-scale legged robot design and fabrication with compliant mechanisms. Our approach makes use of a process that integrates compliant flexure hinges and rigid links to form parallel kinematic structures through the folding of flat-fabricated sheets of articulated parts. Using screw theory, we propose the formulation of an equivalent mechanism compliance for a class of parallel mechanisms, and we use that compliance to evaluate a scalar performance metric based on the strain energy stored in a mechanism subjected to an arbitrary load. Results from the model are supported by experimental measurements of a representative mechanism. With the insight gained from the kinematic mechanism design analysis, we propose and demonstrate compliant designs for two six-legged robots comprising the robotic, autonomous, crawling hexapod (RoACH) family of robots. RoACH is a two degree of freedom, 2.4 gram, 3 cm long robot capable of untethered, sustained, steerable locomotion. RoACH’s successor, Dy- naRoach, is 10 cm long, has one actuated degree of freedom and is capable of running speeds of up to 1.4 m/s. DynaRoACH employs compliant legs to help enable dynamic running and maneuvering and is three orders of magnitude more efficient than its milli-scale predecessor. We experimentally demonstrate the feasibility of a biologically-inspired approach to turning control and dynamic maneuvering by adjusting leg stiffness. While the result agrees qualitatively with predictions from existing reduced order models, initial data suggest the full 3-dimensional dynamics play an important role in six-legged turning.

1 To my grandmother, Millicent Murdock.

i Contents

Contents ii

List of Figures vi

List of Tables ix

Acknowledgementsx

I Introduction and Background1

1 Introduction2 1.1 Design and Fabrication Challenges for Milliscale Mobile Robots...... 3 1.2 Biological Inspiration...... 5 1.3 Contributions...... 6

2 Previous Work8 2.1 Simple Models for Legged Locomotion...... 8 2.1.1 Rigid Inverted Pendulum...... 9 2.1.2 The Spring-Loaded Inverted Pendulum (SLIP)...... 9 2.1.3 The Lateral Leg Spring (LLS)...... 10 2.2 Legged millirobots...... 11 2.2.1 Compliant robots...... 11 2.2.2 MEMS-based robots...... 13 2.3 Dynamic Locomotion...... 14 2.3.1 Hopping...... 14

ii 2.3.2 Multi-legged Meso-Scale Robots...... 15

II Design Principles and Tools for Folded Millirobots 18

3 A Strain Energy-Based Fitness Metric for Compliant Mechanism Designs 19 3.1 Fabrication of Millirobotic Compliant Mechanisms...... 19 3.1.1 SCM - Smart Composite Microstructures...... 19 3.1.2 Existing Approaches to Modeling SCM Mechanisms...... 21 3.1.3 The Pseudo-Rigid-Body model...... 21 3.1.4 Limitations of the PRBM...... 23 3.2 Preliminaries...... 23 3.2.1 Rigid Transformations in 3-Space...... 23 3.2.2 Wrenches...... 24 3.2.3 Transformation of the Compliance Matrix...... 24 3.3 Modeling Mechanism Stiffness...... 25 3.3.1 Compliance of a Small Beam Element...... 25 3.3.2 Compliance Matrix of a Linear Beam...... 26 3.3.3 Compliance Transformation for Individual Flexures...... 29 3.4 Application to Compliant Mechanism Design...... 30 3.4.1 An Energy-Based Design Metric...... 30 3.4.2 Design Evaluation and Analysis...... 31 3.5 Experimental Results...... 36 3.6 Concluding Remarks...... 38

III Legged Robot Designs and Experiments 40

4 Design and Prototyping of RoACH, a 2.4 Gram Hexapedal Robot 41 4.1 Prototyping the RoACH Hexapod...... 41 4.1.1 Kinematic Sketch...... 42 4.1.2 Prototype Fabrication...... 43 4.1.3 Prototype Testing - First Steps...... 44 4.2 Design and Fabrication at Scale...... 46

iii 4.2.1 Fabrication Process...... 46 4.2.2 Mechanical Design...... 47 4.3 Electronics Design...... 58 4.3.1 Control Board...... 58 4.3.2 Power and Processing...... 59 4.3.3 Wireless Communication...... 59 4.3.4 SMA Power Supply...... 60 4.3.5 SMA Drive Method...... 61 4.3.6 Communication/Software...... 63 4.4 Experimental Results...... 63 4.4.1 Turning Control...... 65 4.4.2 Energetics...... 65 4.5 Concluding Remarks...... 66

5 DynaRoACH: Design and Dynamic Maneuverability of a Hexapedal Runner 68 5.1 Robot Design...... 68 5.1.1 Kinematic Transmission Design...... 69 5.1.2 Leg Design...... 73 5.1.3 Power, Communication, and Control Hardware...... 73 5.2 Models for Dynamic Turning in Hexapods...... 75 5.3 Experimental Setup...... 76 5.3.1 Substrate...... 76 5.3.2 Leg Stiffness...... 77 5.3.3 Camera Calibration...... 77 5.3.4 Tracking Geometry...... 78 5.3.5 State Estimation...... 78 5.4 Results...... 79 5.4.1 Running Speed...... 79 5.4.2 Cost of Transport...... 79 5.4.3 Turning by Stiffness Control...... 80 5.5 Adjustable Stiffness Leg Design...... 82 5.6 Discussion...... 82

iv IV Conclusions and Appendices 86

6 Conclusions 87 6.1 Future Work...... 87 6.1.1 Improved Design Tools...... 87 6.1.2 Integration of Electronics into Robot Structure...... 88 6.1.3 Dynamic Characterization and System ID...... 88 6.1.4 Dynamic Climbing...... 88 6.1.5 Robots as Platforms for Testing Biological Hypotheses...... 89

A A Scaled Prototyping Process for Folded Millirobots 90 A.1 Process Overview...... 90 A.1.1 The Prototyping Process...... 90 A.1.2 Flexure Design...... 93 A.1.3 Compensation for Laser Dimensions...... 93 A.1.4 Developing a Scaling Law...... 95 A.2 Comparison of Scaled Prototype to RoACH Millirobot...... 95

Bibliography 98 References...... 98

v List of Figures

1.1 The UC Berkeley Micromechanical Flying Insect...... 5

2.1 Rigid, inverted pendulum walking...... 9 2.2 A schematic of the spring-loaded inverted pendulum (SLIP) model...... 10 2.3 The Lateral Leg Spring (LLS) model of locomotion...... 11 2.4 HEXBUG...... 12 2.5 Previous smart composite microstructures robots...... 12 2.6 Previous work in legged millirobots using MEMS...... 13 2.7 Dynamic hopping robots...... 15 2.8 Dynamic, meso-scale, multi-legged running robots...... 16

3.1 A step-by-step illustration of the MFI SCM process...... 20 3.2 Schematic illustration of an SCM flexure hinge...... 21 3.3 (a) Continuum and (b) Pseudo-Rigid-Body models of a beam flexure...... 21 3.4 (a) Continuum and (b) Pseudo-Rigid-Body models of a small length flexure arrangement...... 22 3.5 A small element in a beam subjected to a wrench...... 25 3.6 Schematic illustration of an SCM flexure hinge...... 27 3.7 A CAD model of a hexapedal crawler...... 31 3.8 A compliant Sarrus linkage...... 32 3.9 Performance of a Sarrus linkage consisting of two kinematic chains...... 33 3.10 Performance variation in a Sarrus linkage comprised of three kinematic chains... 34 3.11 Performance variation in a Sarrus linkage comprised of four kinematic chains... 35 3.12 Experimental Sarrus linkage topologies and geometries...... 36

vi 3.13 Experimental setup for measuring off-axis displacements...... 37 3.14 Comparison of theory to experimental measurements of Sarrus linkages...... 38

4.1 Diagram of the alternating tripod gait at four extreme positions...... 42 4.2 Prototype crawler kinematics showing two independent degrees of freedom.... 43 4.3 Assembly progression of a scaled prototype...... 43 4.4 Time sequence of steps from the robot prototype’s alternating tripod gait...... 45

4.5 Robot prototype climbing a 30◦ incline...... 46 4.6 A step-by-step illustration of RoACH’s SCM process...... 47 4.7 A palette of flat SCM parts...... 48 4.8 Frequency response (above) and output power (below) for a 25 µm SMA...... 49 4.9 Frequency response (above) and output power (below) for a 37.5 µm SMA..... 50 4.10 A CAD model of the robot with body-fixed basis...... 51 4.11 Crawler kinematics depicting two independent degrees of freedom...... 52 4.12 Output from a simulation of the robot kinematics...... 53 4.13 SMA actuator routing...... 54 4.14 A kinematic model of a fourbar linkage with compliance...... 56 4.15 Force displacement curve for a biased fourbar linkage...... 56 4.16 Fourbar model and actual hip fourbar on robot...... 57 4.17 Kinematic model of half of a Sarrus linkage with compliance...... 57 4.18 Force displacement curve for a Sarrus linkage...... 58 4.19 Top and bottom of the control electronics PC board...... 60 4.20 Miniature power supply efficiency for 1 MHz and 2 MHz switching frequencies.. 61 4.21 Miniature boost converter pc board...... 62 4.22 Transistor drive scheme for SMA...... 62 4.23 Current drive scheme for SMA actuators...... 63 4.24 Electronics block diagram...... 64 4.25 The 2.4 g RoACH standing over a US quarter...... 64 4.26 Composited frame captures showing controlled turning...... 66 4.27 Cost of transport data from [33]...... 67

5.1 The DynaRoACH robot...... 69 5.2 Ideal robot kinematics...... 70

vii 5.3 Schematic of fourbar kinematics...... 71 5.4 Schematic of slider crank kinematics...... 72 5.5 3D foot trajectories...... 74 5.6 A schematic of the lateral leg spring (LLS) model...... 75 5.7 LLS turning parameters...... 76 5.8 The robot chassis with rigid tracking geometry and retro-reﬂective spheres attached. 78 5.9 Forward speed vs. stride frequency...... 79 5.10 Metabolic cost of transport vs. stride frequency for the DynaRoACH robot..... 80 5.11 Summary data from all turning and straight running trials...... 81 5.12 Turning rate vs. stride frequency...... 81 5.13 A CAD depiction of the compliant semi-circular leg with shape memory alloy wires 82 5.14 Experimental measurements of dynamic leg stiffness...... 83 5.15 Roll and pitch angle vs. phase for characteristic trials...... 84 5.16 Angular orientation for the robot during dynamic turns...... 84 5.17 Position data for the robot during a sequence of dynamic turns...... 85

A.1 Steps for prototyping a hexapod crawling robot...... 91 A.2 A step-by-step illustration of the posterboard link and hinge fabrication process.. 92 A.3 Effect of workpiece orientation on ﬂexure geometry...... 94 A.4 Designed and actual ﬂexure cut lengths...... 94 A.5 Flexure jamming angle vs measured cut length...... 95 A.6 The scaled prototype and actual structure...... 97

viii List of Tables

3.1 Sarrus linkage geometric parameters...... 36

4.1 Map of actuator states to gait positions shown in Fig. 4.1...... 42 4.2 Mass breakdown for the entire robot...... 59 4.3 Flexinol sizes and resistance per length...... 60 4.4 Summary of the physical parameters for RoACH...... 65 4.5 A comparison of similarly sized legged robots...... 65

5.1 Robot physical parameters...... 70 5.2 Roll and pitch angle statistics for n=14 straight-running trials...... 85 5.3 Roll and pitch angle statistics for n=31 turning trials...... 85

A.1 Comparison of a 2 posterboard prototype to S-2 glass ﬁber composite using Eqn. × A.2...... 96

ix Acknowledgements

First and foremost I would like to thank my advisor, Professor Ron Fearing. His support, guidance, and knowledge over these past six years cannot be measured. It’s truly been a pleasure to learn from him and to have him as a mentor. When I joined the Biomimetic Millisystems lab, I was stunned by the talent, patience, and willingness to help demonstrated by the graduate students and post-docs alike. Rick Groff was incredibly helpful with everything from prelim exam problems to paper edits. Srinath Avadhanula demonstrated an uncanny willingness to teach me new concepts and suggest solutions to research problems I was facing. Rob Wood gave me an invaluable education in the art of trash-talking on the foosball table - a skill that continues to come in handy today. Carmel Majidi has helped me out of more than one continuum modeling jam and continues to be a good friend to this day. Finally, enormous thanks go to Erik Steltz who is responsible for my barely passable circuits and embedded systems literacy. In recent years, I’ve had the good fortune to be surrounded by great labmates who’ve provided stimulating discussions and welcome distractions. In particular, Jongho Lee, Stan Baek, Paul Birkmeyer, and Kevin Peterson have demonstrated admirable willingness to shirk all academic responsibilities in the service of a good game of ping pong. To Bryan Schubert and Andrew Gillies I offer sincere thanks for a willingness to engage any problem and provide valuable insight even in areas unrelated to their specialties. Xiao-Yu Fu deserves many thanks for being the best “post- undergrad” assistant a graduate student could ask for. Without his help over the past six months, this dissertation wouldn’t be finished right now. I’d also like to thank all of the undergraduates I’ve had the opportunity to work with over the years - Minerva Pillai, Dianne Weeks, Andrew Parra, and Maribel Hudson. They each helped teach me a lot about how to improve as a mentor, and I’ve thoroughly enjoyed keeping in touch with them and learning about their latest academic exploits. For educating me in the mysterious ways of biology I must thank Professors Bob Full, Robert Dudley, and Mimi Koehl. For providing a truly memorable experience in my computer graphics coursework and later valuable feedback as one of my committee members I thank Carlo Sequin.´ His passion for engineering and intense curiosity are infectious, and I am humbled by his geometric problem-solving intuition. As my time in Berkeley draws to a close, I can reflect fondly on many pleasant experiences of first making new friends through shared intellectual interests and then watching those friendships grow over coffee and lunch and the occasional beer. Among the many friends who’ve made Berke- ley a supremely enjoyable experience Athulan Vijayaraghavan, Alan Van Pelt, Yonatan Munk, and Maria Jose´ Fernandez stand out. Most recently, I’ve also had the pleasure of working with and getting to know Sam Burden. I look forward to continuing to collaborate with him and possibly even one day having the opportunity to “wheeeze the joooo hooose.” Finally, I would like to thank my family. My sister, Melissa, and my mom, Kathy, have been unwavering in their support of me. They’ve always expressed a sincere interest in my work and genuine desire for me to succeed. Most of all, I want thank my fiancee,´ Robin, for all her care and

x loving support throughout my graduate career. It wasn’t always clear that keeping me sane was a task that could be accomplished, but she took up the mantle and did it with aplomb.

xi Part I

Introduction and Background

1 Chapter 1

Introduction

Any intelligent fool can make things bigger, more complex, and more violent. It takes a touch of genius – and a lot of courage – to move in the opposite direction.

-Albert Einstein

Though much maligned, the humble cockroach is one of nature’s most remarkable examples of high performance legged locomotion. The American cockroach, Periplaneta americana, for example, is capable of achieving running speeds of up to 50 body lengths per second [33][27]. It can traverse rough and smooth terrains alike with little or no modiﬁcation to the train of neural signals activating its leg muscles [82], and it shows remarkable robustness to disturbances ranging from forced vertical running transitions, to falls from heights of many body lengths, to loss of entire limbs [26]. Finally, anyone who’s ever attempted to squash one of these bugs can’t help but appreciate its unparalleled toughness. From a robotics perspective, a system with the characteristics of a cockroach represents the “holy grail” of terrestrial locomotion. With the ability to nimbly run over loose or rugged terrain and climb vertical surfaces quickly, a very small legged robot could prove invaluable for penetrat- ing tiny openings in rubble to search for survivors after an earthquake. Many nearly unnoticeable bug-like robots equipped with sensors and radios could be quickly dispersed over a region to form a reconﬁgurable, mobile sensor network that could have applications like surveillance, or environ- mental and structural monitoring in sensitive areas. Finally, large numbers of small, inexpensive legged robots, each programmed with a simple set of local rules, could be used to study emergent, or swarm behavior on dynamic platforms. However, the design and development of such small scale legged robots is still in its nascency, and challenging engineering problems abound. In some cases, biological insight is helpful - organisms often utilize strategies that are remarkably simple, yet elegant, to solve problems of locomotion like gait coordination, robustness to disturbances, and maneuverability. In other cases, we must draw upon our knowledge of engineered systems to solve problems inherent to building robots at the small scale like the fabrication and actuation of mechanisms. The broad goal of this thesis is to combine engineering insights, with biological inspiration where appropriate, to design and build legged robots at or near the size scale of some of nature’s most incredible locomotion

2 “machines” - cockroaches. While the design approach as well as the robot designs themselves in this thesis have yielded a certain amount of success with respect to that goal, the work presented here represents just the ﬁrst steps in the direction of truly capable legged robots at the insect scale.

1.1 Design and Fabrication Challenges for Milliscale Mobile Robots

The quest to build a capable legged robot at the scale of a cockroach is ambitious. Organisms themselves are “assembled” from a complex hierarchy of materials comprising cells, tissues, and organs at levels of increasing complexity and organization. Furthermore, evolution has had mil- lions of years to tune their fitness, constantly refining and improving upon previous generations of designs. Nonetheless, we are drawn to the challenge of engineering very small robots at the millimeter scale (millirobots) because they have great potential as mobile robotic platforms. Millirobots’ small size provides myriad advantages compared to their larger robot counterparts: it enables them to operate in environments where large robots would be impractical or impossible; at a few dollars per robot, their low cost means they could be produced in large quantities inexpensively; large numbers of them can be networked to form highly mobile and robust sensor and communication networks [54]. By distributing the challenges of mobility, computation, and sensing, it is possible to gain overall system robustness and added capability compared to a single, large mobile robot. In addition to those features mentioned above, small robots have another unique advantage over large robots: their surface area to volume (and therefore mass) ratio is relatively high. One important ramification of this fact is that fluid drag force (which scales proportionally to surface area) begins to dominate inertial forces associated with the robot’s mass at speeds that are relevant to everyday operation. This can create striking physical robustness on par with some of nature’s most resilient arthropods by, for example, enabling a small legged robot to survive falls at its terminal velocity [15]. However, designing capable milliscale robotic systems means first addressing some engineering challenges that are unique to the small scale. As mechanical systems shrink in size, the nature of the forces acting on the system components begins to change. Inertial forces scale with robot volume or proportionally to L3 where L is some characteristic dimension. Surfaces forces like electrostatic attraction, Van der Waals, and capillary scale with surface area or proportionally to L2 [30]. Thus the ratio of the surfaces forces to inertial forces scales as:

Surface Forces 1 (1.1) Inertial Forces ∝ L Minimizing the mechanical interaction of surfaces becomes relatively more important as size decreases. Additional constraints imposed by small size mean milliscale robotic system designs necessarily become tightly coupled and the design must be considered holistically rather than as a aggregate of subsystems.

3 A major contribution that attempts to address this issue of tight system coupling at the milliscale has come from the UC Berkeley Biomimetic Millisystems Lab in the form of the smart composite microstructures (SCM) paradigm [91]. At its core, SCM is a planar fabrication process that integrates rigid, fiber-reinforced polymer composites with compliant homogeneous polymer films to create structures that can be folded up into functional, articulated compliant mechanisms whose mobility is enabled by elastic deformation of flexural elements. In addition to enabling the creation of milliscale articulated mechanisms, SCM structures can be directly integrated with a variety of “smart” material actuators [94][73][45] as well as wiring [74][23]. The first successful millirobotic structure to make use of SCM was the Berkeley micromechanical flying insect (MFI) [90]. Robots produced by the SCM process typically consist of complex interconnections of parallel kinematic structures whose design is driven by a need to minimize the number of actuators required to control their motion. In previous millirobotic designs [95][9][73] a large amount of work has gone into devising mechanical transmissions capable of converting typically small motions of actuators [94] into large, coordinated outputs of wings or legs. In the purely kinematic design phase, these mechanisms may be modeled as consisting of ideal rigid links connected by ideal revolute joints with zero stiffness. This is a convenient abstraction from a design perspective as it allows for the design of motion curves and the optimization of mechanisms using traditional geometric approaches to kinematics. In this abstraction an ideal lower pair like a revolute joint enables one degree of freedom and enforces five degrees of constraint. In practice however, SCM produces joints whose mobility is a result of compliance, and those joints have compliance in all directions. The design process for a mechanism comprising these joints and links is therefore one of enabling compliance in desired directions to produce degrees of freedom and reducing it in others to create degrees of constraint. Successful mechanism and robot designs making use of SCM often require many iterations and rely on the designer’s intuition about material behavior and process limitations to tune their performance. To improve the design process, in Chapter3 we present one approach to modeling and analyzing SCM mechanisms that makes use of screw theory combined with a strain energy-based performance metric from structural design to assess the fitness of SCM mechanism designs. The SCM process (described in more detail in Chapter3) is a powerful tool enabling the fabrication of very small robots, but it also comes with some drawbacks. As with any small-scale fabrication technology, design and assembly can be time-consuming and error prone. For example, a typical millirobot design like the UC Berkeley micromechanical flying insect (MFI) [10] shown in Fig. 1.1 can have feature dimensions as small as 100µm, requires precise integration of very small actuators [93], and is fabricated from pre-impregnated composite materials [91] which require long cure times. Fabrication of an SCM assembly using pre-impregnated carbon fiber consists of a multi-step process of alternately laser micromachining and curing layers which can take one to two days depending on the complexity of the part. Small feature sizes typically necessitate the use of an optical microscope and delicate and careful hand assembly techniques to ensure kinematic integrity of the completed structure. This hand assembly process is error-prone, and errors in assembly can be difficult or sometimes impossible to recover from. Actuators must be precisely aligned to enable the delivery of appropriate forces to the structure and the environment. The overall complexity of these processes makes them time consuming and significantly reduces the rate at which new designs and improvements to existing designs can be implemented. Typically, from

4 the start of cutting to the assembly of a ﬁnished artifact can take a minimum of three days. For these reasons, the small-scale process is not well suited for use in the early stages of robotic design during which many ideas and designs will be explored, but most will be discarded. In our initial attempts to conceive of and build new legged millirobot designs, we were motivated by the above limitations to develop a prototyping process that closely mimics SCM, but eliminates the most time-consuming step of curing composites, as well as painstaking hand assembly at the milliscale. The result is the scaled prototyping process described in Appendix A that uses commonly available, inexpensive materials to reduce the time it takes from days to hours to fabricate and assemble a robot prototype. And, because SCM is a planar approach to creating compliant mechanisms, designs resulting from the scaled prototyping process can be scaled down trivially according to simple geometric scaling rules. Chapter4 provides a validation of the prototyping approach by demonstrating a functional prototype of a six-legged millirobot at a 2X scale. In the same chapter we demonstrate the feasibility of scaling the prototype to produce an actual autonomous, untethered, legged millirobot at the size scale of a cockroach and capable of sustained 5 steerable walking.

8mm

Figure 1.2: MFI prototype from 2006. Figure 1.1: An example robot built using the smart composite microstructures (SCM) paradigm - a version of the UC Berkeley micromechanical ﬂying insect from 2006. the determination of this is addressed in this thesis.

1.2Another Biological secondary problemInspiration is that the two wings are independent resonant systems. Con- sidering the structure is assembled by hand, the resonant frequencies of the two wings are very SCM provides an engineering solution for small scale fabrication, but in order for these highly constrained mobile millirobotic systems to be capable of autonomous locomotion, they must be difficultpowered to match. and The controlled. resulting Actuation, structure power, has and to controloperate at at the a milli-scale drive frequency are very averag challenginged from the research problems which require very different approaches from the control systems employed in resonant frequencies of the two wings; this results in less than optimal wing flapping angle for both 5 wings. Finally, the airframe for the structure is not adequate to completely isolate effects from one wing to the other. When trying to control wings independently, some of the control action from one wing influences the other due to airframe vibrations, which makes control very difficult.

1.5 Contributions and Outline

The contributions of this work are as follows:

Power density framework for millirobots - Most millirobots in the literature have limited • mobility because they were not optimized in terms of power density of the robot. Chapter 2

presents a framework for designing millirobots in terms of power density of the power plant of

the robot. Several actuators, batteries, and drive electronics are compared to show the strength

and weakness of each actuation method and how to choose a technology given a locomotion macro-scale robotic systems. Weight and size constraints restrict the number and size of actuators, making uncontrolled degrees of freedom unavoidable. Limitations on power density impose a trade-off between locomotion and computation; achieving rapid motion restricts the allowable complexity of control algorithms. Taken together, these constraints necessitate increased reliance on mechanical control for these robots. This tight coupling requires the robot designer to consider simultaneously, and in an integrated fashion, issues of fabrication, actuation, power, and control. In this arena, turning to the biological world for inspiration enables us to extract important principles that can inform our choices in mechanical design, actuation, and control. Although they can be highly complex and sophisticated, there are notable cases in which they use simpliﬁed control strategies that still enable high performance locomotion. The cockroach, Blaberus discoidalis, can traverse very large obstacles without changing neural activation patterns to its hind leg muscles responsible for propulsion [82], all the while maintaining a roughly constant gait pattern. With regard to maneuverability, the same cockroach can generate asymmetric force patterns that enable it to make fast, dynamic turns with very small modiﬁcations of its leg stiffness and gait kinematics [52]. Reduced order dynamic models that attempt to explain these unique system properties like self-stabilization and dynamic turning have recently been the subject of much study [16][17][52][76][71] and are discussed in more detail in Chapter 2. The goal of the work presented in Chapters 4 and 5 is to couple some of these underlying principles from biology with the design and fabrication approaches described in Chapter 3 to enable capable legged robots despite the challenging constraints imposed by working at the small scale.

1.3 Contributions

This work makes the following contributions:

Fitness metric for evaluating compliant, folded robot designs - An approach to analyzing • the compliant mechanisms resulting from the SCM process that makes use of an equivalent mechanism stiffness formulation combined with an energy-based metric from structural analysis. The result is a scalar metric that is useful for characterizing the ﬁtness of a folded compliant mechanism design from the perspective of its resistance to off-axis loads. We also demonstrate a qualitative agreement between the theoretical model and experimental results.

RoACH - the smallest untethered, autonomous crawling robot to date - The design and • complete kinematic model of a 2.4 gram, untethered hexapedal robot using a modiﬁed SCM process, integrated shape memory alloy wire actuators, and onboard power and control. The robot is steerable and capable of sustained quasi-static locomotion. To the author’s knowledge, this robot is the smallest, untethered, power-autonomous legged robot to date.

DynaRoACH - The design and kinematic model of a 24 gram, dynamic hexapod driven by • a single DC motor. The robot is capable of dynamic turning by adjusting leg stiffness and, as such, represents the ﬁrst effort to use a biologically inspired control approach to dynamic turning. We demonstrate a distributed compliance, viscoelastic leg for high-speed running

6 enabling running speeds of up to 14 body lengths per second at stride frequencies of up to 20 Hz. Measurements of the dynamic stiffness of the C-shaped leg are provided, and we demonstrate the feasibility of an actuated leg design using shape memory alloy wire to adjust leg stiffness as well as length.

A scaled prototyping process for folded millirobots - A low-cost, fast, scaled version of • the Smart Composite Microstructures process that enables complete design and assembly of a compliant robotic structure from inexpensive materials in just a few hours is presented in AppendixA. The prototyping process enables the robot designer to iterate over a large design space quickly and has already enabled the design of viable new robots in the Biomimetic Millisystems Lab.

7 Chapter 2

Previous Work

It’s not hard to understand the considerable interest among engineers in using lessons from biology to inform the design of mobile robotic systems. Nature provides us with a wealth of exemplary displays of high performance locomotion — cockroaches running over rough terrain [82][4], geckos climbing a variety of vertical surfaces and even running upside down [51], and fruit ﬂies executing 90◦ turns called saccades in less than 50 ms [31]. Observing and modeling these systems can reveal important principles that robotics engineers can take advantage of to improve performance, enable novel behaviors, or simply create mobile robots at scales heretofore impossible. In this chapter, we ﬁrst provide a brief review of models for legged locomotion that have emerged from the study of organisms. These reduced-order models allow us to discern key principles at work in legged systems, but are simple enough to admit parametric analysis in most cases. Following the exposition of the reduced order dynamic models, we present a review of a few of the most recent milli- and meso-scale legged robots, some of which are capable of dynamic locomotion.

2.1 Simple Models for Legged Locomotion

The study of legged locomotion from a biological perspective can be daunting — an enormous number of legged animals exist all with varying morphologies, body sizes, and preferred natural environment. However, biologists have worked to show that some aspects of legged locomotion follow single functions irrespective of body size, leg morphology, or preferred operating environment. For example, two-, four-, six-, and eight-legged animals all demonstrate qualitatively similar ground reaction force patterns [32][33]; animals that vary in body mass over 5 orders of magnitude all use about the same amount of energy to move one kilogram of their body mass one meter [40] (though larger animals do so more efﬁciently [84]); even despite differing leg numbers many animals use equivalent gaits [6]. Given these similarities across species, morphology, body size, and operating environment, biologists have constructed reduced-order models that attempt to capture features of legged locomotion that are common to a variety of organisms.

8 2.1.1 Rigid Inverted Pendulum

The simplest model for legged locomotion is the rigid, inverted pendulum model of walking [5][20] restricted to the sagittal plane. Walking is modeled as a point mass atop a massless rigid rod as depicted in Fig. 2.1. In this model, kinetic energy of the mass is exchanged for gravitational potential at the apex of the stride and then back into kinetic energy at the end of the stride. If no double support phase (both legs in contact with the ground) exists, some kinetic energy must be dissipated on foot impact. With a double support phase, the leading leg does negative work on the center of mass while the trailing leg does positive work in redirecting the velocity vector of the center of mass. Stride parameters like length and width (out of plane) can help to explain the energetic cost of walking using this model [58]. And, despite the fact the model is essentially bipedal, it is possible to treat multi-legged locomotion by considering groups of legs (three legs in a six-legged animal, four in an eight-legged animal, etc.) as single functional legs. SS DS SS V+

V-

Figure 2.1: Walking modeled as a rigid inverted pendulum. The SS portions are single support in which only one leg is in contact, and DS is double support (light grey shaded areas), during which time both legs are in contact with the ground. V − is the velocity of the center of mass just before touchdown and V + is the velocity of the center of mass just after double support ends.

2.1.2 The Spring-Loaded Inverted Pendulum (SLIP)

The SLIP model [16][17] depicted in Fig. 2.2 is an extension of the inverted pendulum to include vertical compliance of a massless leg. Adding leg compliance makes the model much more dynamically rich while still restricting the configuration to a tractable number of system parameters that determine performance. Because of the leg compliance, it is possible for the leg to lose contact with the ground at the end of a stride given the appropriate initial conditions. This behavior makes SLIP a hybrid dynamical system, that is the time evolution of the state of the system is governed by two distinct sets of differential equations. One set governs the stance phase when the leg is in contact with the ground (beginning at a touchdown event) and the other governs the ballistic or flight phase (beginning at a liftoff event). The system switches between the two continuous vector fields based on events of liftoff (LO) and touchdown (TD). The SLIP model assumes that the touchdown angle of the leg is reset to the same initial value every stride.

9 Despite being a conservative system, SLIP has been demonstrated to have stable limit cycles [36] using both linear and air spring laws for the spring-loaded legs. More recently, Seipel and Holmes demonstrated the clock-torqued SLIP (CT-SLIP), an actuated version of SLIP that includes damping and a clock-driven torque action at the hip that more accurately predicts performance and stability seen in robots and natural systems with SLIP-like dynamics [78].

landing velocity

reaction gravitational force force

Figure 2.2: A schematic of the spring-loaded inverted pendulum (SLIP) model

2.1.3 The Lateral Leg Spring (LLS)

The LLS is a reduced-order dynamical model of legged locomotion in the horizontal plane [76][77]. The LLS, a schematic of which is shown in Fig. 2.3, was motivated by obser- vation of side-to-side bouncing of cockroaches in the horizontal plane and therefore models legged locomotion in the horizontal plane as a single mass with two legs sprawled out to either side of the body. Like SLIP, the touchdown angle of each leg is prescribed at the beginning of each stride. Motion of the LLS model can be parameterized by the initial leg length, leg touchdown angle, leg stiffness, and the distance of the center of pressure from the center of mass along the major axis of the body. Also like SLIP, despite the fact that it’s a conservative dynamical system, LLS shows self-stabilizing properties with respect to heading (velocity vector of the center of mass) and body orientation.

LLS and Maneuverability

In addition to the interesting self-stabilizing properties of the LLS model, Proctor and Holmes also demonstrated the possibility of inducing turns simply by momentarily destabilizing the system [71]. Changes to model parameters like leg stiffness and location of the center of pressure with respect to the center of mass can shift the magnitude of one of the eigenvalues of the Jacobian of the stride map to temporarily destabilize the system and cause translational kinetic energy to be exchanged for rotational, thereby inducing a turn. Our experimental results in Chapter 6 provide support for the hypothesis that turning is possible with small differential adjustments to leg stiffness.

10 STEERING BY TRANSIENT DESTABILIZATION 275 precisely straight we seta ¯ = 1, ¯b = +1 (dashed CoM path), and to reverse angular momentum − and turn right we seta ¯ = 2, ¯b = +1.5 (dot-dashed CoM path). These respectively give A =1.618, A = 0 and A = 0.615 for− the turning strides, after whicha ¯ and ¯b revert to the values 0.5, 1.0 for which A =0.−25 and the paths rapidly straighten. In the limited context of the peg-leg− model, moving the TD point forward along the body mimics extension of the outside front leg observed in insect turning [24], and it provides our ﬁrst evidence that foot position can modulate a turn.

3. THE LATERAL LEG SPRING MODEL In this model a single effective leg represents the support tripod of front, rear and contralateral middle legs that many insects, including cockroaches, employ during normal running [3, 4]. A rigid body, again with mass m and moment of inertia I, is equipped with two massless elastic legs that pivot freely at a point P on the centerline, a distance d from the CoM (d>0 denoting toward the front of the body). A left (resp. right) stride (stance phase) is initiated at TD by placing the relaxed leg, with length l0, at an angle +β (resp. β) relative to the body axis and allowing the foot F to pivot freely about the resulting position,− which remains fixed until the leg spring is again uncompressed, at which point LO occurs and the right (resp. left) stance phase immediately ensues (a 50% duty cycle). The center of pressure (CoP) d is fixed throughout each stride. See Fig. 8(a), and refer to [7] for further details.

Figure 2.3: A schematic from [76] illustrating the relevant parameters of the lateral leg spring (LLS) model for locomotion in the horizontal plane Fig. 8. The LLS model showing the body and leg geometry (a) and the coordinate system for the left stance phase, pivoting at F (b). That for right stance is obtained analogously, with TD angle β, as described in the text. − 2.2 Legged millirobots

Interest in legged millirobots3.1. has begun Equations to take off of in Motion recent years. and Perhaps Stride one to of the Stride best Map indicatorsTo describe of this increased dynamics interest is during in the commercial a stride availability we adopt of milliscale a polar robotic coordinate toys. The systems based at the stance Hexbug [49] shown in Fig. 2.4 is a 5 cm long, 15.6 g hexapedal legged robot driven by a single foot.motor but Since still capable the leg of walking is elastic, (and turning) the system in reverse. now It’s equipped has three with a microphone degrees of and freedom (two translation and onebump sensors rotation), (at the base described of the antennae) by the that enable generalized it to sense obstacles coordinates and noises (ζ, inψ order) that to locate the CoM r(t), and theinitiate body an escape angle maneuverθ. Fig. that consists 8(b) of shows turning the while coordinate backing up. The convention Hexbug’s kinematics for left stance, with ψ increasing counter-clockwise;prevent it from being able tofor traverse right anything stance butψ theincreases smoothest clockwise,surfaces (several but piecesθ is of taken paper increasing counter-clockwise throughout.stacked are capable To of tripping accommodate the robot up), this but it’s we nevertheless preface anθ impressivewith the robotic factor toy. ( 1)n where left and right strides correspond to n even and odd respectively [7]. In terms of these coordinates− the kinetic and potential energies2.2.1 Compliant are given robots in Eqs. (14–15): 1 ˙2 2 ˙2 1 ˙2 The ﬁrst attempt at a legged millirobot using theT SCM= 2 processm(ζ came+ ζ fromψ )+ Sahai2etI al.θ [,73]. The (14) proposed robot depicted in Fig. 2.5a design integrates power, communication, control electronics, 2 2 n and battery into a 35 mm long, 3 g package.V = V The(η) leg, η drive= mechanismζ + d consists+2ζd ofsin a novel (ψ kinematic( 1) θ), (15) design that couples two piezoelectric bending actuators to the six legs - one to each tripod− - through− 2 n def a parallel arrangement of slider crank linkages.LF = Althoughmζ ψ˙ !+( this robot1) didIθ˙ not= walkpψ duep toθ excessive= const. (16) serial compliance, it was the ﬁrst legged robot to integrate all systems− required for locomotion± with the SCMREGULAR fabrication AND process. CHAOTIC DYNAMICS Vol. 13 No. 4 2008

11 2 A.T. Baisch and R.J. Wood

capable of running at speeds up to 20 body lengths per second (Blaberus discoidalis) and even 40 body lengths per second (Periplaneta Americana) [8]. Cockroaches are known to maintain high speed ambulation across rough, uneven terrain [21] and provide further design inspiration with their ability to climb sheer vertical and even inverted surfaces [17]. Furthermore, they have efﬁcient navigational aids such as antennae that allow obstacle avoidance even in low lighting [13]. These aspects of insect morphology enable robust locomotion, and we strive to translate the underlying principles into designs for an insect-scale ambulatory robot. Roboticists have already proven the importance of using nature as a design guide, as evidenced by legged robots that are more agile than their wheeled counterparts [19],[15],[16]. There has been work to create climbing robots for rough [11] or smooth [12] surfaces, and even efforts to integrate insect-like proximity sensors for navigation along walls [14], [4]. Others have attempted to use MEMS fabrication to create silicon insect-scale ambulatory robots [6], [27], [5]. Our goal is to combine many of these functions into a single autonomous millimeter-scale robot. This work focuses on the design and fabrication of thoracic and leg mechanics, actuation, and electronics required to achieve such a goal. Fur- Figure 2.4: HEXBUGthermore weR : examine A commercially potential modes available of attachment 16 g, hexapedal as a preliminary millirobot study onwith scal- auditory and touch sensinging capabilities. vertical surfaces with a microrobot. We conclude with an overview of remaining research topics towards the creation of a fully autonomous ambulatory robotic insect.

2 Robot Design

Dry Adhesion Mechanism Leg Transmission

Proximal Distal ‘knee’ Actuator Mount Joint (a) (b) Fig. 1 Notional design of the thorax and legs of the Harvard Ambulatory Microrobot (left) and Figure 2.5: Previousinitial prototype work (right) in legged millirobots using the smart composite microstructures (SCM) fabrication paradigm. (a) A fully integrated, six-legged, piezoelectrically actuated 3 g crawling robot [73]. (b) AThe mobile design six-legged of the Harvard millirobot Ambulatory with Microrobot eight independent (HAMR), degrees presented of in freedom, Fig. 1, actuated with shape memorywas motivated alloy [12 by]. several factors, including available fabrication techniques, and the necessary body mechanics to achieve gaits inspired by insects. Important constraints in designing a microrobot are brought about by the fabrication techniques and materials that may be used. A solution will be discussed in Section 3, however our

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G$=9,&!DI!T./&,$0&*($*=!/1:3)1+!')*2&:1*'&!)*!4556/&+&!;$(#! More recently, work from the Harvard Microrobotics Laboratory has explored using shape -&33:V9/!/)%$($)*!B1E!4556/&+&!$*!-&33:V9/!/)%$($)*!;$(#!%$.!3&=%>B-E! memory alloys in a legged millirobot. Fig. 2.5b shows the Harvard Ambulatory Microrobot Q$3$')*!/1:3)1+>B'E!]1:3)1+!)*!(#&!3&=%>B+E!K7L@!Y))0V$*!2$&;!)C!(#&! (HAMR), a 17 mm x 22 mm robot fabricated using the SCM process with ten integrated SMA C&1(9,& )* (#& /1:3)1+ 9%&+ C), (,1'H$*= ! actuators to control its gait. The robot has been demonstrated at actuation frequencies of up to 8&! ')*+9'(&+! /1:3)1+! ')*2&:1*'&! &./&,$0&*(%! 9*+&,! 1 Hz, and is capable of performing zero radius turns. In addition to actuated locomotion, HAMR 21,$)9%! ')*+$($)*%! )C! 4556/&+&! 3&=! 2&3)'$($&%7! 6*! (#&%&! is capable of supporting 100 times its own weight while attached to a vertical surface using micro- &./&,$0&*(%>! ;&! (,1'H&+! 0$',)%'13&! C&1(9,&%! )*! ()/! )C! (#&! claws. The robot is powered and controlled by offboard electronics through an attached tether. /1:3)1+!9%$*=!1!K7L@!0$',)%')/&!3&*%!1*+!M1-2$&;N!6O4P! 01'#$*&! 2$%$)*! ()! +&(&,0$*&! (#&$,! 3)'1($)*7! "#&%&! C&1(9,&%! G$=9,&!ZI!4556/&+&!0$',),)-)(!+$%/31'&0&*(!/&,!':'3&!C),!+$CC&,&*(! 1,&! C1-,$'1(&+! 9%$*=! QR6! S56T! )*! (#&! %$3$')*! /1:3)1+7! 6*! C,$'($)*!0)+&3%7!")/I!5)-)(!+$%/31'&0&*(!/&,!%(&/[!\)(()0I!M&=!%3$/7! (#$%! '1%&>! 133! 3&=%! 1,&! 1'(91(&+! ()! 0)2&! ;$(#! (#&! %10&! 2.2.2 MEMS-based robots "#&!+,$C(!)-%&,2&+!$*!@J!$%!1!*90&,$'13!1,($C1'(!1*+!$%!)2&,')0&!-:!(#&! &.&'9($)*!)C!1!%9-%&F9&*(!%(&/!$*!(#&!=1$(!%&F9&*'&7! 2&3)'$(:7!8&!&./&,$0&*(&+!;$(#!')0-$*1($)*%!)C!C,&F9&*':! ! 1*+! 10/3$(9+&! ,&%93($*=! $*! %$0$31,! 2&3)'$($&%! 1*+! ! &./&,$0&*(133:! 0&1%9,&+! (#&! ,&%93($*=! +$%/31'&0&*(! )C! (#&! ! ,)-)(!/3)((&+!$*!G$=9,&!<7!8&!*)($'&+!(#1(!(#&!%$+&;1:!+,$C(! ! B13)*=!UE!;#&*!(#&!10/3$(9+&VC,&F9&*':! /,)+9'(%! )*! -)(#! ! %$+&%! )C! 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(#&! %(&&,$*=!$%!%$0$31,7! grated solar power and control electronics [43]. (c) The thermally actuated ARRIpede [68], shown 4556/&+&! ')*'&/(7! 8&! 9%&+! 1*! $*2&,(&+! /,)()(:/&! ;$(#! !1*+!01*$/931(&+!1!?!'0!@!?!'0!@!A00! ')*(,)33$*=!(#&!2&3)'$(:!)C!(#&!,)-)(!$%!')9*(&,$*(9$($2&>!%$*'&! %$3$')*! +$&! 1%! 1! /1:3)1+! BC$=9,&! DE! /);&,$*=! $(! &.(&,*133:7! &3&'(,)(#&,013!1'(91(),%!1,&!*)*3$*&1,7!X);&2&,>!$(!01:!-&!1! "#$%! /1:3)1+! ;&$=#%! 1//,).$01(&3:! (#&! %10&! 1%! 1! ,)-)(! =))+! 1//,).$01($)*! -&'19%&! -:! ':'3$*=! (#&! =1$(! 1(! Some attempts at locomotion at%9//),($*=! the millimeter $(%! );*! scale ;&$=#(! have also B?=E7! been "#&! made 0$',)',1;3&,! using traditional 3&=%! 1,&! C,&F9&*'$&%! '3)%&! ),! #$=#&,! ()! (#&! (#&,013! -1*+;$+(#! MEMS processes. Ebefors silicon microrobot1'(91(&+! 9*+&,! [28] is 21,:$*=! just 15 mm 10/ x3$(9+&%! 5 mm, weighs 1*+! C,&F9&*'$&%! 80 mg, and is ()! BDKXYE>! (#&! 1((1$*1-3&! 2&3)'$($&%! 1,&! /,)/),($)*13! capable of walking at speeds of up to')*(,)3! 6 mm/s. (#&! Shown /)%$($)*! in Fig. )C! 2.6a (#&!, the /1:3)1+7! robot uses "#&! thermal /,)+9'(! actuators )C! (#&! consisting of soft-baked polyimide deposited10/3$(9+&! on ;$(#! silicon (#&! inC,&F9&*':! a standard )C! MEMS 1'(91($)*! process. =$2&%!1!0&1%9,&! Geometry of the leg joints is such that heating)C! of (#&! the 3&=!polyimide 2&3)'$(:! creates 1*+! $%! a thermal 9%&+! 1%! expansion (#&! ')*(,)3! that /1,10&(&,! causes the ()! joint to deﬂect and the leg to engage%(&&,! the surface. (#&! ,)-)(7! One limitation G),! &.10/3&>! of this ;#&*! approach 3))H$*=! is that ();1,+%! the thermal (#&! actuators require a considerable amount+$,&'($)*!)C!0)($)*>!')*%$+&,!(#&!'1%&!;#&*!(#&!3&=%!)*!(#&! of power (approximately 1.1 W at the fastest actuation speeds) thus, on-board power was deemed,$=#(!1,&!1'(91(&+!1(!#$=#&,!2&3)'$(:!(#1*!(#&!3&=%!)*!(#&!3&C(>! infeasible, so the robot operates on a wire tether. Hollar et al. used a silicon-on-insulator$7&7I! (SOI) process to create a two-legged robot 8.5 mm x 4 mm with a mass of only 10.2 mg [43]. The robot is remarkable in that the mechanical system consisting of compliant joints, rigid silicon legs, and electrostatic inchworm motors, the solar power supply, and a sequencer chip for controlling leg motion are all integrated and on board. In initial tests, the robot was capable of generating 3 mm of sideways motion. The fully integrated 835 system is depicted in Fig. 2.6b Authorized licensed use limited to: Univ of Calif Berkeley. Downloaded on April 16,2010 at 06:03:35 UTC from IEEE Xplore. Restrictions apply. While not strictly legged, Bergbreiter and Pister have developed a MEMS-based robot intended to locomote not by walking or running, but rather by jumping [13]. The design makes use of a

13 novel arrangement of electrostatic inchworm actuators to stretch a micro-rubberband fabricated from molded elastomer. A release mechanism enables very fast release of stored elastic energy ultimately intended to launch the robot into the air. A macro view of the whole robot can be seen in Fig. 2.7c. The most recent legged millirobot built using traditional MEMS processes is the 1 cm x 1 cm, 3.8 g ARRIpede from Murthy et al. [68]. Like Ebefors’ robot, the ARRIpede in Fig. 2.6c also uses electrothermal actuators. Unlike Ebefors robot, ARRIpede’s actuators are prismatic. With a power consumption of approximately 750 mW, the ARRIpede is capable of walking at 1.5 mm/s. However, the ARRIpede includes a power electronics and battery backpack that enables it to walk untethered and accounts for most of the weight of the robot.

2.3 Dynamic Locomotion

With the exception of Bergbreiter’s hopper, all of the robots mentioned above use a quasi-static approach to locomotion. Dynamic locomotion is quite a bit more complex, and the transition from quasi-static locomotion modes like walking to dynamic running or hopping is generally discrete and abrupt. Dynamic locomotion often requires the coordination multiple limbs (or, at the very least, one limb and one body), entails impulsive ground interaction, and can involve complex 6 DOF motions coupling angular and translational degrees of freedom to induce surprising behaviors. However, it is exactly this complicated behavior that makes these dynamic systems a rich area for robotics research.

2.3.1 Hopping

In some ways, hopping represents a logical starting point for the study of dynamic legged locomotion. It’s inherently dynamic as it often involves the sudden release of stored energy and appropriate redirection or dissipation of that energy on landing. In fact, as is demonstrated in the SLIP abstraction of dynamic running, in the sagittal plane running can be reduced to hopping with some forward speed. Some of the first successful examples of dynamic legged locomotion were products of research from the Leg Lab at MIT under the direction of Raibert. There, they designed and built the 3D single leg hopping robot shown in Fig. 2.7a. The monopod was capable of balancing itself and following prescribed paths on a flat surface [72]. In addition to the 3D monopedal hopper, Hodgins and Raibert designed, built, and programmed a bipedal robot capable of performing flips [42]. The flipping biped used a simple primarily feedforward, open loop control strategy with some event synchronization based on sensory feedback. Many of Raibert’s robots relied on a pneumatic air supply and offboard power and control, so most operated on a tether. More recently, in the space of hopping robots, Kovacˇ has developed a 14 g, 18 cm spherical hopping robot driven by a single DC motor [57]. The robot uses a cam to wind a spring and rapidly release the elastic energy to hop. The robot depicted in Fig. 2.7b is capable of hopping almost 3.5

14 296 Auton Robot (2010) 28: 295–306

The “Sandia hopper” (A) (Weiss 2001) is a 2.5 kg system that jumps by using a piston driven combustion cham- ber mounted inside a spherical, plastic shell. After landing it remains in its stable position and can rotate its center of mass in order to change its orientation and jump again. The 128 stair hopping “scout robot” (B) (Stoeter et al. 2002) is a wheeled system of 200 g that jumps using a spring that is coiled around its body. Locomotion in a given direction is ensured by turning the robot prior to jumping. The “minimalist jumping robot” (C) (Burdick and Fiorini 2003) has a mass of 1.3 kg and can, after landing, extend its structure actively and upright for the next jump. The change of direction happens by rotating on its foot. The Jollbot (D) (Ar- mour et al. 2007) is a 465 g robot that can actively squeeze its spherical structure in order to upright after landing and can change the direction of its jump by rotating its center of (a) (b) gravity around its(c) main axis. The “rescue robot” (E) (Tsuk- agoshi et al. 2005) has a mass of 2 kg and can, due to its Figuresemispheres5.7: Fully on theIntegrated sides ofMo theckup. wheels, upright on land- Figure 2.7: Dynamic hoppingFig. robots 1 Prototype at with the a mass macro-, of 14 g. The cage meso-, has a height and of 18 cm milli-scales.ing, then turn using(a) itsRaibert’s wheels very self- similar to the “scout balancing, hopping 3D monopedaland allowsrobot the robot to [ upright72]. itself (b) after 18 landing, cm steer tall and jump steerable again robot” hopper (B) and capable jump using a of pneumatic hopping drive. Table 1 sum- The energy stored by the inchworm motor was calculated by assuming that a to heights of more than three times it’s body length and reorienting upon landing. [57]. (c)A milliscale hopper fabricated using MEMS technologymaxim integratedum force of with225 µN elastomerswas used to stretc forh the energymicro rubb storageer band. Because the micro [13]. rubber band is assembled “pre-strained,” energy is also stored in the micro rubber band

during assembly as described in Section 3.2.4. A load force of 100 µN from pre-straining

was estimated by assembling a similarly fabricated micro rubber band into a force gauge

Fig. 2 Overview of existing jumping robotselsewhere that areon ablethe tochip, per- and(Cappro) “minimalistximately Jumping3 nJ w Robot”,as stored v2 (Burdickdue to andpre-straining Fiorini 2003dur) ing form repetitive steered jumps. (A) “Sandia hopper” (Weiss 2001) (pic- (reprinted by permission of SAGE), (D) Jollbot (Armour et al. 2007) times its own height and rightingture is and courtesy orienting of Sandia National itself Laboratories), after (B landing,) “stair-Hopping making(reprinted it by possible permission of IOP), for (E) the “rescue robot robot” (Tsukagoshi et al. assembly. Therefore, it was estimated that the inchworm stored an additional 1.9 nJ of to “steer” itself toward a goal.scout robot” (Stoeter et al. 2002) (reprinted by permission of IEEE), 2005) (reprinted by permission of IEEE) energy before slipping for a total of 4.9 nJ stored and quickly released. The new inchworm Table 1 Steerable jumping robot comparison

Name Massmotors Max.designed Jumpin Chapter Jump4 will Jumpincrease heightthis stored Jump heightenergy to Uprightingthe tens of SteeringmicroJoules 2.3.2 Multi-legged Meso-Scale(Fig. 2) Robots [g] size height distance per mass per size principle principle needed[cm]for tens [cm]of centi [cm]meters jumps. [cm/g] [–] (Fig. 4) (Fig. 5)

(A) Sandia robot 2500 20 300 300 0.12 15 (C) (B) In the past ten years, a variety(Weiss 2001 of robot) designs inspired5.4.3 Fi bynal runningRobot Mo arthropodsckup like cockroaches have emerged. One of the most(B) Stair-Hopping successful scout robot six-legged 200 11 robots 30 to 20 run dynamically 0.15 2. is7 the RHex (C) (A) (Stoeter et al. 2002) Finally, a mechanical mockup of what the ﬁnal robot might look like was assembled robot [75][8], a version of which(C) Minimalist is shown jumping robot in Fig. 1300 2.8a 15. RHex 90 is 200 a 7 Kg, 0.07 six-legged 6 robotwith (A) (C) (Burdick and Fiorini 2003) compliant legs that rotate continuously and are eachand drivenis shown byin aFigur singlee 5.7. motor.This mockup Asgiv such,es some it’spersp capableective on the ﬁnal size of the (D) Jollbot 465 30 21.8 0 0.06 1.4 (B) (B) (Armour et al. 2007) of mounting obstacles higher than its hip height androb hasot as evenwell as demonstratedbetter ideas on how to stair-climbingassemble the separate [66comp]onen as ts. (E) Rescue robot 2300 15 80 unknown 0.03 5.3 (C) (A) well as stair descent [19]. (Tsukagoshi et al. 2005)

EPFL jumping robotTM 14 18 62 46 4.17 3.4 (C) (D) Shortly after RHex was developed,(presented here) the Whegs family of legged robots was also released, the smallest of which is Mini-Whegs [67]. The Mini-Whegs in Fig. 2.8c is a 150 g quadruped that also uses a hybrid wheel-leg design similar to RHex, but unlike RHex it uses a single motor to drive all four legs, with servo motors to provide steering. At 9 cm long, Mini-Whegs is capable of running 10 body lengths per second and surmounting obstacles approximately 1.5 times the radius of its wheel-legs. The Sprawl family of robots from Stanford, are also directly inspired by the kinematics and dynamics of cockroach locomotion. Sprawlita [21] is a six-legged robot that’s fabricated using a pro-

15 Highly Mobile and Robust Small Quadruped Robots Jeremy M. Morrey, Bram Lambrecht, Andrew D. Horchler, Roy E. Ritzmann, Roger D. Quinn

Case Western Reserve University, Cleveland, Ohio, U.S.A. http://biorobots.cwru.edu, [email protected] (a) (b)

Abstract This paper describes novel highly mobile small robots called “Mini-Whegs” that can run and jump (see video). They are derived from our larger Whegs series of robots, which benefit from abstracted cockroach locomotion principles. Key to their success are the three spoked appendages, called “whegs,” which combine the speed and simplicity of wheels with the climbing mobility of legs. To be more compact than the larger Whegs vehicles, Mini-Whegs uses four whegs in an alternating diagonal gait. These 9 cm long robots can run at sustained speeds of over 10 body lengths per second and climb obstacles that are taller than their leg length. They can run forward 1cm and backward, on either side. Their robust construction allows them to tumble down a flight of stairs with no damage and carry a payload equal to twice their weight. A jumping mechanism has also been developed that Figure 1. Photograph showing(c) relative sizes of the Mini- (d) enables Mini-Whegs to surmount much larger obstacles, Whegs IV robot and a Blaberus gigantius cockroach. such as stair steps. Figure 2.8:they offer Previous much of workan advantage in dynamic, over wheels meso-scale, because at multi-legged running robots. (a) RHex: a 7 Kg, 1.0 Introduction six-leggedthis small robot scale capable it is difficult of to dynamic implement a running, modern track stair-climbing, and a variety of biologically-inspired suspension [1]. Fukui et al. [6] developed a small hexapod Highly mobile small vehicles, sometimes calledgaits [75robot]. (b)that usesiSprawl: piezoelectric a 16 actuators cm longto run in six-legged a tripod runner with cable actuation driven by a single micro-robots, are better suited for certain missions motor.than gait. iSprawl The vehicle can runis limited at speeds to relatively up flatto 15 surfaces body lengths per second [55]. (c) Mini-WhegsTM: a larger vehicles. For example, they can aid in search and because its legs have short ranges of motion. rescue because their diminutive size enables them towheel-leg fit It hybrid is difficult quadruped for small robots that to canmove jumpthrough andreal- run 10 body lengths per second [67]. (d) DASH: a into tight spaces, such as those found in rubble and16 in g six-leggedworld terrain runner simply because driven of bythe arelative single size motor, of the capable of running at 15 body lengths per second caves. As another example, a group of small robots obstacles they must overcome. Therefore, it is particularly provide robustness through redundancy for remoteand withstandingimportant for impacts small robots at to its use terminal efficient locomotory velocity [15]. missions such as extraterrestrial exploration. Mobile small appendages. For a given vehicle size, legs promise the robots are also appropriate for insect inspired research greatest mobility because they enable discontinuous because their scale is similar to that of the insect models. contact with the substrate, which is important for uneven A variety of robots similar in size to the vehicles terrain. Insects are excellent examples of highly mobile, described in this paper have been developed, but the legged vehicles and therefore a robot designer would be majority of them have limited mobility. For example, well advised to draw inspiration from them. Khepera robots have a 5cm wheelbase and 1.4cm Biological inspiration can be implemented in varying diameter wheels [9]. They can move only on very smooth, degrees from the direct to the abstracted [13]. The direct flat surfaces. The robots in the Alice series use relatively approach sometimes requires that new technologies be large wheels as compared to the size of the robots for developed, whereas abstracted locomotion principles can improved mobility, but they still suffer from the often be implemented using current technology. limitations of wheels on complex terrain [3]. Scout also The design of Sprawlita [4] was inspired by the uses relatively large wheels for its size and some versions cockroach. At 16cm long, it is larger than the other robots of it have a separate mechanism that enables it to jump up described here. It is a hexapod that uses a combination of 16 a stair or use expanding wheels to overcome larger servo motors and air cylinders. Its top speed of 4.5 body obstacles [5]. Millibots use tracks but it is not clear that lengths per second is very fast as compared to existing cess called shaped deposition manufacturing (SDM) that makes it possible to integrate viscoelastic compliant polymers with rigid, structural polymers as well as enable embedding of sensors and actuators in the actual robot body itself [64]. Sprawlita used pneumatic actuators as primary power sources to drive prismatic legs that are attached to the hips with compliant, viscoelastic hinges that helped to stabilize the robot motion by providing tuned damping. Because of the pneumatic actuation, Sprawlita has not been able to run autonomously. However, a later Sprawl family robot based on very similar kinematics, iSprawl (shown in Fig. 2.8b), is capable of autonomous locomotion as it uses a single DC motor to drive all six legs through a slider crank mechanism connected to push- pull cables that operate the prismatic legs [55]. The key insight of iSprawl was the introduction of series compliance to the leg design that enabled it to go from running 5 body lengths per second to running 15 body lengths per second. The most recent addition to the current lot of biologically inspired six-legged running robots is from UC Berkeley. At just 16 g, and 10 cm long, the Dynamic Autonomous Sprawled Hexapod (DASH) [15] shown in Fig. 2.8d is one of the lightest and fastest legged robots for its size. It’s fabricated using the scaled SCM process [44], is driven by a single DC motor, and is capable of speeds up to 15 body lengths, or 1.5 m per second. It’s unique design and construction also enable it to withstand impacts at its terminal velocity.

17 Part II

Design Principles and Tools for Folded Millirobots

18 Chapter 3

A Strain Energy-Based Fitness Metric for Compliant Mechanism Designs

Compliant mechanisms are a useful class of mechanisms because they attain their mobility from the elastic deformation of compliant members. The term “compliant mechanism” has been broadly used in the literature to describe both continuum mechanisms in which large scale deformation can be distributed throughout the body [98][86] and so-called discrete or lumped compliance models in which compliance is concentrated at nodes which connect members that are generally assumed to be rigid [48][47]. Because they have no mechanical parts whose surfaces move with respect to each other, they are free from backlash and friction and require no lubrication, making them advantageous for use at the microscale. These advantages have led to their adoption in a variety of robotic applications from microscale silicon-based robots [28][43], to high resolu- tion mesoscale precision positioning devices [11][24], to milliscale mobile robots with microscale feature sizes [89][45] as well as the scaled prototypes discussed in the previous chapter.

3.1 Fabrication of Millirobotic Compliant Mechanisms

3.1.1 SCM - Smart Composite Microstructures

Robot fabrication at the millimeter scale is inherently difficult with conventional mechanical components like continuous revolute joints, bearings, and motors because as component size is reduced, surface forces begin to dominate Newtonian forces [87]. Some MEMS research has been successful in creating robotic mechanisms like hinged joints [70], folded structures [97], and mobile microrobots [28][43][13]. However, MEMS processes require an expensive infrastructure, are restricted to a specific set of materials compatible with existing microfabrication processes, and the design cycle can be very long. SCM was designed expressly for the fabrication of milli-scale articulated components with micro-scale features. It’s a process that integrates laser micromachined composite fiber laminates with polymer films to create mechanisms consisting of rigid composite links and compliant poly-

19 mer hinges as shown in Fig. 3.2. The process as it has been implemented for the fabrication of previous robots in the Biomimetic Millisystems Lab is shown in Fig. 3.1 and the steps are outlined below.

1. The process begins with a sheet of uncured, pre-impregnated composite ﬁber laminate.

2. Gaps are laser micro-machined into the composite ﬁber.

3. A polymer ﬁlm (typically polyimide) is then placed on top of the laminate.

4. The polymer sheet is cured to the composite ﬁber.

5. The cured composite and polymer are released and aligned onto another composite ﬁber layer.

6. The parts are cured.

7. The resulting ﬂat structure is released.

Figure 3.1: A step-by-step illustration of the SCM link and hinge fabrication process as implemented in the fabrication of the Berkeley micromechanical ﬂying insect (MFI).

Fig. 3.2 is an illustration of a flexure hinge created using the SCM process. A layer of polymer film sandwiched between two cured sheets of carbon fiber provide a large range-of-motion compliant flexural joint that is devoid of friction and backlash and as a result does not wear. Entire structures consisting of rigid joints as well as flexible joints can be fabricated in this fashion. The structure can be folded from its flat position into a fully 3D structure - flexure hinges can be glued in place to form permanent rigid joints or left free to serve as compliant hinges for motion transmission as the design requires. The process has the advantages that an entire articulated robotic structure can be fabricated in a single flat “piece,” and, unlike MEMS technology, it can quickly and easily produce truly 3D structures.

20 Figure 3.2: Schematic illustration of an SCM ﬂexure hinge

3.1.2 Existing Approaches to Modeling SCM Mechanisms

The most typical approach to modeling of compliant mechanisms is the ﬁnite element method (FEM) in which the body is subdivided into mesh elements and a large system of equations relating the tractions to the displacements for each elastic element are solved numerically. This method can be very powerful for optimizing or predicting the behavior of an existing topology [41], but its complexity and lack of analytical predictive power make it a cumbersome tool to apply during initial design or prototyping stages.

3.1.3 The Pseudo-Rigid-Body model

        

  (a) (b)

Figure 3.3: (a) Continuum and (b) Pseudo-Rigid-Body models of a beam ﬂexure

In the early stages of prototyping, designs tend to change very quickly, making time-consuming finite element analysis infeasible. At theses stages, to simplify the design process and provide a closed form approximation for the behavior of loaded flexures it is helpful to employ an approximate model of flexure stiffness. One such model that is useful for this purpose is the pseudo-rigid- body model (PRBM). The PRBM enables a mechanism designer to replace a continuum model of a flexural element with a lumped parameter model consisting of two rigid links connected by a linear torsional spring at some distance along the length of the original element as depicted in Fig. 3.3.

21 " !

Kθ !

(a) (b)

Figure 3.4: (a) Continuum and (b) Pseudo-Rigid-Body models of a small length ﬂexure arrange- mentWednesday, May 12, 2010

If the flexural element connects two comparatively rigid bodies and can be considered “small- length,” it’s possible to reduce the flexure joint to an ideal pin joint connecting the two rigid members in parallel with a linear torsional spring as shown in Fig. 3.4. More specifically, [47] defines flexure pivots as “small-length” if they meet the following criteria:

Llink >> Lflexure (3.1)

(EI)flexure << (EI)link (3.2)

In Eqn. 3.1 Llink is the length of a comparatively rigid link to which one end of the flexure is attached, and in Eqn. 3.2 the flexural stiffness, (EI)link is the flexural stiffness of that link. Characteristic link dimensions in millirobots are on the order of multiple millimeters while flexure lengths are typically 100-200µm, so this requirement is generally satisfied by the designs we are interested in creating. If a flexure can be considered small length, the expression for its rotational stiffness is reduced to the following:

EI K = (3.3) θ L In Eqn. 3.3, E is the flexural modulus of the material from which the flexure hinge is constructed, I is the area moment of inertia, and L is the length of the flexure. This approximation is very useful in the early stages of design because it enables the designer to quickly estimate forces and torques for a given mechanism design. However, because the model is planar and accounts for only a single degree of freedom, it cannot give any useful information regarding mechanism behavior in the presence of off-axis loads. This limitation has motivated the analysis in the rest of this chapter wherein SCM mechanisms are treated as spatial mechanisms which are designed to produce desired degrees of freedom as well as degrees of constraint.

22 3.1.4 Limitations of the PRBM

The PRBM was developed in an effort to simplify the compliant mechanism design process and bridge the gap between rigid mechanism design and compliant mechanism design. Most of the time, the PRBM is sufficient to model and analyze flexure mechanisms, and has been used with success in a variety of applications such as compliant robot leg design [34]. However, in practice, in our complicated kinematic designs, flexures experience loading and the consequent displacements in as many six axes. For larger scale devices or mechanisms where the applied and reaction loads are consistent and predictable this limitation of the PRBM may not be significant. But, for small scale mechanisms incorporated into millirobots where actuation involves the amplification of small strains or actuator output motions [94], the consideration of additional off-axis loads and displacements becomes increasingly important for avoiding the loss of actuator displacement to serial compliance in the structure. Put another way, our compliant mechanisms are designed to provide degrees of freedom and degrees of constraint. We desire that the mechanisms are most mobile in the directions enabled by their degrees of freedom and least mobile in the constrained directions determined by their degrees of constraint. In cases where we are willing to assume small deflections, we may explicitly incorporate these multi-axis effects (in closed form) into our model in order to better study the off-axis behavior of a mechanism. In this chapter, we address the problem of improving the design of a single degree of freedom (DOF) compliant mechanism consisting of rigid links connected by simple flat flexure hinges experiencing small deflections. Specifically, we are interested in improving the various off- axis stiffnesses of such mechanisms while preserving the output motion in the desired direction. The example mechanism to which we apply our analysis here is the compliant version of a simple 1 DOF exact straight line linkage (shown in Fig. 3.8) known as the Sarrus linkage after its inventor, French mathematician Pierre Fred´ eric´ Sarrus.

3.2 Preliminaries

3.2.1 Rigid Transformations in 3-Space

Rigid motions in space can be represented as screws, or translations along a line plus a rotation about that line. A screw, sT = [ω, v], is an element of the Lie algebra of SE(3). As such, it is transformed actively by the action of an element from the adjoint representation of the group:

s0 = Ad(g)s (3.4)

where Ad(g) is a 6x6 matrix with the following form:

R 0 Ad(g) = (3.5) SR R

23 For the remainder of this chapter we’ll refer to this rigid transformation as T. Further decom- posing T enables us to write it as:

I 0 R 0 T = (3.6) S(r) I · 0 R In Eqn. 3.6, the matrix S(r) is the cross-product matrix deﬁned as follows:

0 r r − z y S(r) = rz 0 rx (3.7)  r r −0  − y x   So, we see that R is the rotation matrix that deﬁnes the portion of the transform that consists of a rotation of the screw, and S accounts for the translation portion of the transformation with r being the vector from the initial position of the rigid body to its transformed position. The active transformation also has a corresponding passive change of coordinates that’s given by the inverse of the transformation:

T 1 R 0 ω s = T− s = (3.8) RTSRT v −

3.2.2 Wrenches

A wrench can be thought of as a generalized force. It generalizes forces and moments in the same way the screw generalizes translations and rotations into a six-dimensional vector. Wrenches are elements of the dual to the Lie Algebra and so transform according to the inverse transpose of T:

T R SR M w = T− w = (3.9) 0 R f

3.2.3 Transformation of the Compliance Matrix

The compliance matrix of a body is a generalization of the concept of one-dimensional compliance into multiple dimensions. The compliance matrix relates applied forces and moments (wrenches) to linear and angular displacements (screws). More precisely, the compliance matrix is a linear operator that transforms wrenches into screws. In this chapter we will use rigid coordinate transformations of the compliance matrix extensively, so we note that combining Eqns. 3.8 and 3.9 yields the equation determining how the compliance matrix transforms under a coordinate change:

1 T C = T− CT− (3.10)

24 3.3 Modeling Mechanism Stiffness

In this section, we derive the expression for the compliance matrix of a beam by ﬁrst starting from a small beam element. This derivation was done in [79], but it is repeated here for complete- ness. Then, using the compliance matrix resulting from that derivation, we describe a process for referring the compliance matrix of each ﬂexure in an SCM mechanism to a common coordinate system. This compliance transformation will enable us to express an “equivalent” mechanism compliance in a common output coordinate system. The focus on a rectangular cross-section beam of uniform thickness is motivated by the compliant structures we are currently capable of producing using the SCM process.

3.3.1 Compliance of a Small Beam Element Δµ z y x w w

Δθy

Figure 3.5: A small element in a beam subjected to a wrench, w.

Fig. 3.5 shows a small element from a beam subjected to a wrench, w. The beam element is assumed to be in equilibrium. The beam element has a ﬁnite compliance that transforms the applied wrench into a screw displacement expressed as the difference of two screws:

Cw = ∆s (3.11)

In Eqn. 3.11 C is the compliance matrix for the small beam element. In this analysis we will assume small deﬂections, so that ∆µ ∆s where s is the arc length of the beam. This is the ≈ ∂2y same assumption that enables curvature to be approximated as ∂x2 in the classic analysis of a beam proceeding directly from the Euler-Bernoulli equation. Additionally, we assume that cross-sections of the beam remain perpendicular to the neutral axis even in the deformed state. Starting with the axial force component of the wrench, we have the following relation describing the displacement of the beam due to an applied force along the beam axis:

1 ∆x F = (3.12) x EA ∆µ 25 Where E is the modulus of the material and A is the cross sectional area of the beam element. For a moment about the x axis, the angular displacement is dependent on the size and shape of the cross section. Here we are only considering a beam with a rectangular cross section of constant thickness. However, an approximate expression relating the moment about the x axis and the angular displacement in the y z plane is: − 1 ∆θ M x (3.13) x αG ≈ ∆µ where α is a function of the geometry of the cross-sectional area. Formulas for α for a rectangular cross-section beam can be found in [35]. A moment applied about the y axis results in pure bending of the beam element and gives the familiar equation:

1 ∆θy My = (3.14) EIz ∆µ

3 where Iz is the area moment of inertia deﬁned for the beam geometry shown in Fig. 3.6 as bh /12. Similarly, a moment applied about the z axis also results in pure bending:

1 ∆θz Mz = (3.15) EIy ∆µ

3 where Iy is the area moment of inertia about the y axis: b h/12. We choose to neglect the effect of shear in the beam, so displacements due to forces in the y and z directions are ignored. Assembling Eqns. 3.12- 3.15 into matrix form gives the following:

Mx ∆θx My ∆θy M  1 ∆θ  c z = z (3.16)  F  ∆µ  ∆x   x     F   ∆y   y     F   ∆z   z        In Eqn. 3.16 c can be considered a compliance per length, ∆µ.

3.3.2 Compliance Matrix of a Linear Beam

We can now use Eqn. 3.16 to compute the compliance matrix for an entire beam. First, however, because c is expressed in a local coordinate frame (located at the center of the element), we

26 b

Figure 3.6: From [39]: A ﬂat ﬂexure depicting local coordinate systems before (x, y, z) and after (x0, y0, z0) deformation. This choice of local coordinate system diagonalizes the compliance matrix.

must ﬁrst transform it to a common coordinate system for the entire beam before we can sum the individual compliance matrices for the differential elements. We parameterize the distance along the x axis from the ﬁxed base of the beam by µ. How we express the location of the coordinate system for each differential element of the beam will depend on our choice of common coordinate system in which to express the whole beam compliance. We choose to take the location of the l origin of our common coordinate system to be located at half the length of the beam, or 2 . The reason for our choice of coordinate system will become clear later. The location of the origin of each coordinate system local to the differential beam elements is given by µ, meaning we must translate the local coordinate system for each beam element by a distance l µ. To do this we construct the following transformation matrix using Eqns. 3.6 and 2 − 3.7:

1 0 0 0 0 0 0 1 0 0 0 0   1 0 0 1 0 0 0 T− = (3.17) 0 0 0 1 0 0   0 0 l µ 0 1 0  2 −  0 l + µ 0 0 0 1  − 2   

27 Combining Eqns. 3.17 and 3.10 generates the following expression for the compliance matrix of the beam as a function of the distance, µ, along the beam:

1 Gα 0 0 0 0 0 µ l 0 1 0 0 0 − 2  EIz EIz  1 l 2µ 0 0 0 − 0 EIy 2EIy c(µ) =  1  (3.18)  0 0 0 AE 0 0   l 2µ (l 2µ)2   0 0 − 0 − 0   2EIy 4EIy  2  l µ l   µ 2 ( 2 )   0 − 0 0 0 −   EIz EIz    Integration of Eqn. 3.18 with respect to µ on the interval [0, l] reveals the reason for our choice l of origin location for the common coordinate system at 2 :

l Gα 0 0 0 0 0 0 l 0 0 0 0 EIz  0 0 l 0 0 0  EIy  0 0 0 l 0 0  (3.19)  AE   0 0 0 0 l3 0   12EIy   3   0 0 0 0 0 l   12EIz    Clearly, our choice of common coordinate system for the compliance of the beam diagonalizes the compliance matrix. This result agrees with the results in [88] and [29]. However, to further verify the compliance matrix, we can take the common coordinate reference to be located at the end of the beam. Elements T54 and T62 in Eqn. 3.17 are replaced by l µ and -l + µ respectively. − Eqn. 3.18 becomes:

1 Gα 0 0 0 0 0 1 µ l 0 0 0 0 − EIz EIz  1 l µ  0 0 0 − 0 EIy EIy c(µ) =  0 0 0 1 0 0  (3.20)  AE  l µ (l µ)2  0 0 − 0 − 0   EIy EIy   µ l (l µ)2   0 − 0 0 0 −   EIz EIz    Integrating 3.20 on the interval [0, l] yields the familiar compliance matrix for a beam in which

28 the cross-coupling between moments and forces applied at the end of the beam is apparent in the off-diagonal terms of C:

l Gα 0 0 0 0 0 0 l 0 0 0 l2 EIz 2EIz  2 −  0 0 l 0 l 0 C = EIy 2EIy (3.21)  0 0 0 l 0 0   AE   0 0 l2 0 l3 0   2EIy 3EIy   2 3   0 l 0 0 0 l   2EIz 3EIz   −  Finally, we note that the above formulation only considers deformation due to bending, torsion, and axial extension - we have not considered shear forces or applied axial loads. A formulation of the diagonal compliance matrix that includes terms representing these effects was presented in [46].

3.3.3 Compliance Transformation for Individual Flexures

Eqn. 3.19 describes the compliance properties of a single flexure, but the mechanisms in which we are interested are typically comprised of multiple flexures connecting sets of rigid links in both serial and parallel configurations. In order to analyze the stiffness properties of an entire mechanism, we must combine the flexure compliance matrix with knowledge of mechanism geometry to generate an equivalent mechanism compliance. Just as we did with individual small beam elements to construct a compliance matrix for the beam, this is begun by transforming the flexure compliance in the local coordinate frame into the output coordinate frame of the mechanism. In general, for any system of springs, elements in parallel are those that experience the same displacement and their stiffnesses can therefore be added to create an equivalent stiffness. Elements in series experience the same load - their compliances add to create a single equivalent compliance. However, this combination is only possible when the stiffnesses or compliances to be added are expressed in a common coordinate system. Therefore, we require that the compliance matrix in (3.19) for each flexure in the mechanism be transformed to a common coordinate system. This transformation refers the compliance of each flexure joint to the base coordinate system (usually attached to the output platform or link of a parallel mechanism), and, in general, will depend on the position and orientation of each flexure relative to the chosen base coordinate system. While the compliance matrix of each flexure explicitly represents a general six axis spring, the spatial transformation introduces the mechanism geometry into the computation of the overall compliance of the mechanism. Here, again, we use 6x6 transformation matrix in Eqn. 3.5 to transform the compliance of an individual flexure in the local coordinate system to the common coordinate system chosen to be at the output of the mechanism. For each flexure in a parallel flexure mechanism the matrices S and R will be functions of the mechanism geometry. For each flexure, we formulate S and R matrices

29 and apply the tranformation in Eqn. 3.10 in order to express the compliance matrix in a common coordinate system. The resulting compliance matrix for the whole mechanism, C0 is the result:

T 1 C0 = T0/i− CiT0/i− (3.22) i X th In Eqn. 3.22, Ci is the compliance matrix of the i ﬂexure in the mechanism and T0/i is the transformation from the ith coordinate frame to the output coordinate frame of the mechanism.

3.4 Application to Compliant Mechanism Design

Using the above formulation, we would like to analyze compliant mechanism designs that result from using the SCM process. In addition to the constraints imposed by the limits of the SCM fabrication process, mobile millirobot design is highly constrained by the size and mass of the electronics and power source needed to power and control actuators. As such, current and past robot designs [96][45] have made extensive use of parallel kinematics in order to appropriately couple and control the desired degrees of freedom while minimizing the number of required actuators.

3.4.1 An Energy-Based Design Metric

With these considerations in mind, we aim to design a compliant mechanism that exhibits maximal compliance in the desired degrees of freedom and maximal stiffness in the undesired degrees of freedom. In order to compare the performance of different designs, it is therefore necessary to adopt a metric by which to judge a mechanism. In this work, we use the ratio of the so-called “mutual potential energy” [80] (also sometimes referred to as “mutual strain energy”) to the total strain energy of the mechanism:

MPE f = (3.23) SE The mutual potential energy is deﬁned as:

MPE = vTKu (3.24)

where v is the displacement vector at the output when only a unit dummy load is applied in the desired output direction, K is the equivalent stiffness of the mechanism expressed in the base or output coordinate system, and u is the displacement vector when the entire input load is applied. Intuitively, we can think of mutual potential energy as the amount of useful work or displacement generated at the output port by an input load. The strain energy of the mechanism is given by:

1 SE = uTKu (3.25) 2

30 Figure 3.7: A CAD model of a hexapedal crawler that uses a Sarrus linkage to connect six fourbar hip joints.

Thus, the performance metric given in Eqn. 3.23 is a ratio of the work output in the desired direction (the useful work) to the work done in deforming the mechanism. The larger this ratio, the more “efﬁcient” the mechanism can be considered, ie. the undesirable degrees of freedom remain stiff while the desirable degrees of freedom are compliant.

3.4.2 Design Evaluation and Analysis

SCM robots’ highly constrained mechanical and kinematic requirements often lead to mechanically complex designs. As a simple example of the usefulness of modeling the full compliance of these SCM structures, we focus on the design of a mechanism that plays a prominent role in the robot described in Chapter4, the Sarrus linkage. Fig. 3.7 shows a model of the robot in which a central Sarrus linkage (shown in blue) can been seen connecting two sets of three hips such that when it contracts the upper legs are adducted while the lower legs are abducted. The Sarrus’ central role in this compliant robotic structure and the complex loads that are applied at the hips as a result of ground reaction forces transferred through the legs make it an interesting case study. Before proceeding with our analysis it is important to note the distinction between topology and geometry as we use the terms here. A mechanism topology refers to the number of links and joints and their connectedness. One can consider an abstraction of a linkage as a graph in which joints are represented by edges of the graph and links are represented as nodes. In this abstraction, the addition of a redundant kinematic chain to a mechanism clearly changes the topology of the linkage as well as its graph representation. Geometry, however, refers to the physical dimensions and conﬁguration of the mechanism. It can include, for example, link lengths, angles between links, and in the case of compliant mechanisms, ﬂexure dimensions. The most common (and minimal) topology of the Sarrus linkage (as pictured in Fig. 3.8)

31 Figure 3.8: A compliant Sarrus linkage shown with its simplest topology and conﬁguration: Two 4 link kinematic chains rigidly joined at their base and output links and oriented at φ = 90◦to each other.

consists of two planar kinematic chains of 4 links connected by three hinges. The two chains are rigidly connected at their base and outputs and are oriented at an angle (usually 90◦) with respect to each other such that the rotation axes of all the joints intersect. Mobility of the chain as predicted by Grubler’s¨ criterion is zero, so it’s important that the rotation axes for the joints be precisely aligned. Using conventional mechanical components, the Sarrus linkage would typically be constructed with pin-type hinges, which ideally provide infinite compliance in rotation about the hinge axis and infinite stiffness in all other directions. However, we are designing a compliant Sarrus linkage in which each hinge has finite compliance/stiffness. Given that fact, we are interested in the topologies and geometries of the Sarrus linkage which, for a given set of forces applied at the output, maximize the metric in (3.23). The minimal topology for the compliant Sarrus consists of two serial kinematic chains connected at some angle, φ, with respect to each other. Conventional Sarrus linkage designs typically set φ = 90◦. In fact, for conventional designs with ideal hinges, any angle other than zero and 180◦ will provide equivalent exact straight line motion constraint. However, for the compliant case, the angle between the two kinematic chains becomes an important design variable when considered in tandem with loading conditions at the output frame. All cases considered below have the geometric parameter θ (depicted in Fig. 3.8) set to 30◦. Additionally, the loading condition considered is the simultaneous application of forces (but no moments) in all three axes.

Case I: Two chain compliant Sarrus

This topology is shown in Fig. 3.8. We are interested in the angle between the two kinematic chains, φ that gives the best performance in the sense of Eqn. 3.23 for forces in the x0, y0, z0 coordinate frame. Fig. 3.9 shows the performance of the two chain Sarrus for the range φ [0, π] ∈

! = 0.5 ! = 1 2 ! = 2 ! = 4

1.5

1 MPE/SE

0.5

0 0 0.5 1 1.5 2 2.5 3 φ (radians) Figure 3.9: Performance of a Sarrus linkage consisting of two kinematic chains connected at an angle, φ. The parameter β is the ratio of the off-axis forces to the force in the direction of highest compliance.

for four different loading conditions. The parameter β represents the ratio of the magnitude of the two off-axis forces (taken to be equal in these examples) to the forces in the direction of highest compliance (the x0 direction, in this case). From the ﬁgure, we see that as the off-axis forces become large in comparison to the force in the compliant direction, a clear optimal orientation angle for the kinematic chains appears at approximately 0.85 radians.

Case II: Three chain compliant Sarrus

From a strictly kinematic perspective, the redundancy introduced by adding a third kinematic chain is unnecessary for the straight line constraint. However, from a compliance perspective, our design intuition is that redundancy should provide an improvement in mechanism performance. Fig. 3.10 shows that, in fact, we do see an improvement in performance. For the case of relatively small off-axis forces (β=0.5) maximum performance is unchanged. However, the mechanism shows a reduced sensitivity to an increase in β at the optimal orientation angle of φ = 0.85 radians. It’s important to note, however, that this relative insensitivity to off-axis forces only occurs at the optimal orientation angle. Away from that angle, performance drops markedly with increased off-axis loading. Additionally, the peak performance predicted for this conﬁguration oc-

! = 0.5 ! = 1 2 ! = 2 ! = 4

1.5

1 MPE/SE

0.5

0 0 0.5 1 1.5 2 2.5 3 φ (radians)

Figure 3.10: Performance variation in a Sarrus linkage comprised of three kinematic chains with two chains connected at angles of φ and -φ with respect to the third. β is the ratio of off-axis forces to forces in the most compliant direction.

π curs to the left of the φ = 2 due to the asymmetric loading conﬁguration. When the force in the y direction is removed, because of the symmetry of the topology and geometry of the mechanism, π the peak performance occurs at the angle φ = 2 .

Case III: Four chain compliant Sarrus

The geometry of this case corresponds to a mirroring of the geometry in case I about the plane z = y. Results from the analysis of this topology are shown in Fig. 3.11. This topology shows the best performance in two senses. First, the maximal performance is relatively insensitive to the orientation angle of the chains (away from the singular φ = 0 and φ = π orientations). Secondly, the performance of the mechanism is relatively insensitive to increases in the off-axis loads over that range of φ.

! = 0.5 ! = 1 2 ! = 2 ! = 4

1.5

1 MPE/SE

0.5

0 0 0.5 1 1.5 2 2.5 3 3.5 φ (radians)

Figure 3.11: Performance variation in a Sarrus linkage comprised of four kinematic chains. This conﬁguration shows the overall best performance encapsulated by a lack of sensitivity to off-axis forces (β) as well as orientation angle (φ).

35 c d e f g a b h Figure 3.12: Sarrus linkage topologies and configurations used to experimentally verify the predictions of the model. a-e are the five configurations of the two chain topology a, φ = 30◦ up to e, φ = 150◦. f-h represent the five configurations for the four chain topology. For f, φ = 30◦(150◦), g, φ = 90◦, and h, φ=60◦(120◦). Only three linkage configurations are necessary for the four chain topology because the 120◦ and 150◦ orientations are reflections of the 60◦ and 30◦ orientations respectively.

Parameter Value b 8mm h 50µm l 600µm llink 20mm θ 10◦ Table 3.1: Sarrus linkage geometric parameters

3.5 Experimental Results

To test the predictions of model, we fabricated 5 different geometric conﬁgurations of two topologies of the Sarrus linkage - the two chain topology and the four chain topology. The mechanisms were constructed using the folding process described in [44] and AppendixA. Table I summarizes the common geometric parameters for all linkages tested. The base of each linkage was grounded and a load was applied to the output link at the free end by a simple mass as depicted in Fig. 3.13. The entire linkage was rotated through several angles to create ratios of off-axis to on-axis loading, β, of 0-2. The linear and angular displacements in the XZ plane were measured optically, and the 3x3 stiffness matrix was estimated from the displacement data using the procedure outlined in [22]. Results comparing the performance of the two topologies are shown in Fig. 3.14.

36 Sarrus linkage 63.5° F = mg

Figure 3.13: Experimental setup used to generate controlled off-axis loading for testing compliant Sarrus linkage performance.

37 2

1.8

1.6

1.4

1.2

MPE/SE 0.8

0.6 2 Chains (Experiment) 0.4 4 Chains (Experiment) 2 Chains (Theory) 0.2 4 Chains (Theory) 0 0 0.5 1 1.5 2 2.5 3 φ (radians) Figure 3.14: Comparison of theory to experimental measurements of two and four chain Sarrus linkages of varying geometries.

For each topology 5 different conﬁgurations (orientation angles of the individual chains) were subjected to an off-axis force in the negative Z direction and a moment about the Y axis as well as an on-axis load in the negative X direction. Fig. 3.14 shows a comparison of the theoretical prediction for the mechanism performance to the experimental data. While the theory and experiment are not an exact match, both bear out the conclusion that the 4 chain Sarrus is preferable to the two chain topology and the peak performance, or maximal off-axis stiffness, for both occurs when π the chains are oriented at 2 radians or 90◦ to each other. The overprediction of the mechanism performance by the model could be due to the fact that the model does not account for types of loading other than bending and normal axial loads. More accurate results may be achievable by integrating, for example, the effect of shear forces.

3.6 Concluding Remarks

By transforming the stiffnesses of simple, constant cross section, beam-like flexures into a common coordinate system we can generate an expression for the generalized stiffness of a compliant mechanism consisting of rigid links interconnected by flat flexures. Using a strain energy-based metric, we can analyze the performance of a given topology and geometry subject to a prescribed set of loads at the output of the mechanism. In this work, this approach has been applied to understanding and improving the off-axis performance of a compliant linkage commonly used in the design of our millirobots - the exact straight line Sarrus linkage. Using the tools described we demonstrated a quantifiable improvement in the performance

38 of the compliant Sarrus linkage by adjusting only the linkage topology. However, we have not optimized in any formal, global sense the topology or geometry of the mechanism. Rather, we have applied a straightforward closed form model to assist in understanding mechanism behavior under general loading conditions. That ability to quickly predict performance improvements of these compliant linkages has the potential to provide additional insights at the early stages of design and reduce the overall design cycle time for SCM millirobots. It is important to note, however, that the results of this approach are dependent not only on the mechanism topology and geometry, but also on the loading conditions imposed at the output of the mechanism. This dependence of the mechanism performance on the loading condition requires the mechanism designer to understand the nature of the output loads associated with the intended application. If those loads are well understood, this approach provides a useful tool for guiding the initial exploration of the design space. Once a promising design has been identiﬁed, more intensive approaches such as ﬁnite element modeling can be applied to improve the precision or performance of the design.

39 Part III

Legged Robot Designs and Experiments

40 Chapter 4

Design and Prototyping of RoACH, a 2.4 Gram Hexapedal Robot

In this chapter, we merge the areas of legged locomotion and millirobotic design to present the Robotic Autonomous Crawling Hexapod (RoACH) a 2.4 g legged millirobot capable of steerable locomotion. The robot makes use of the SCM process for its design and construction and of shape memory alloy wire actuators to provide the mechanical power for locomotion. The design of the robot presented here progresses from a scaled, functional, kinematic prototype using the process described in AppendixA to a true-scale, fully capable untethered millirobot. The prototype is useful for verifying and quickly adjusting the kinematic design, after which point more thorough analysis is performed to complete the design at the milli-scale. This work represents the ﬁrst step in applying the enabling technology of SCM to produce a fully functional, integrated, legged millirobot. As such, we focus primarily on the robot’s mechanical design, the use of the SCM process, and the integration of the actuation, power, and control but do not present a dynamic system model or a control map. The primary contribution of this chapter is demonstrating the viability of legged machines at the milliscale through the integration of novel processes and technologies.

4.1 Prototyping the RoACH Hexapod

Because millirobot fabrication at scale can be a difﬁcult and time-consuming process for the reasons listed in Section 1.1 of Chapter1, and because our proposed design represents a major departure from previous legged robot designs in the Biomimetic Millisystems Lab [73], we have chosen to begin the design process with a 2 scaled prototype of the robot. This approach serves × two purposes: 1) it enables us to verify that the scaled prototyping process in AppendixA works as designed, and 2) it allows us to verify kinematic functionality before committing time and effort to milli-scale fabrication.

41 4.1.1 Kinematic Sketch

Here we will brieﬂy describe a simple kinematic sketch of the scaled prototype of the robot, leaving the bulk of the analysis to be done in the portion of this chapter that deals with the milliscale design. Because the ﬂexure joints that SCM relies on do not admit continuous rotation, it is impossible for the output of a single degree of freedom linkage to produce motion on a closed curve without experiencing a kinematic singularity or some kind of elastic instability such as snap- through buckling. However, effective locomotion that does not rely on stick-slip phenomena or vibrational excitation requires that a “foot” move on just such a closed trajectory. The implication of this limitation of SCM joints is that our kinematic design requires a minimum of two degrees of freedom to enable walking. We would also like to implement a gait that provides static stability to the robot actively (when it’s walking) and passively so that when all of the robot’s actuators are switched off, it is still capable of standing. Looking to nature for biological inspiration reveals the simplest possible gait for a hexapod - the alternating tripod gait [4]. In this gait, the ipsilateral front and hind legs move in phase with the contralateral middle leg, and as long as the center of mass remains over the triangle of support formed by those three contacts, the robot is statically stable. A schematic depiction of the gait is shown in Fig. 4.1.

(a) (b)

Figure 4.1: Diagram of the alternating tripod gait at four extreme positions

Position Actuator 1 Actuator 2 a OFF OFF b ON OFF c ON ON d OFF ON

Table 4.1: Map of actuator states to gait positions shown in Fig. 4.1

42 (a) First degree of freedom - abduction (b) Second degree of freedom - protrac- and adduction (rear view) tion and retraction (side view)

Figure 4.2: Prototype crawler kinematics showing two independent degrees of freedom

(a) Flat prototype parts (b) Component linkages folded (c) Complete folded and assembled prototype

Figure 4.3: Progression of scaled prototype of a hexapedal legged robot from ﬂat parts to folded ﬁnal result.

Fig. 4.2 depicts the kinematic concept for instantiating the two degrees of freedom required to move the foot on a closed trajectory and coupling them appropriately to six legs to create the alternating tripod gait depicted in Fig. 4.1. The concept sketch makes use of parallel kinematics to appropriately couple the legs and generate an alternating tripod gait using just two actuated degrees of freedom.

4.1.2 Prototype Fabrication

A 2X scale prototype of the robot based on the kinematic sketch above was fabricated according the process described in AppendixA. The progression from ﬂat component parts to the folded robot skeleton is shown in Fig. 4.3. There are three primary pieces in the robot prototype design, and they are pictured in Fig. 4.3a. The two similar looking plates are the top and bottom plates of the robot prototype. Each includes the links necessary for folding up three fourbars which altogether form the six hips. Once

43 assembled, each hip is connected to its respective plate (top or bottom) and the middle plate. The hip joint is designed for two degrees of freedom. The ﬁrst of these degrees of freedom is the rotation of the links of the parallel fourbar. With the leg attached, these are the protraction and retraction motions of the leg shown schematically in Fig. 4.2a. The other degree of freedom comes from the fact that the hips of the upper (lower) tripod combined with the middle and upper (lower) plate also form a fourbar linkage. When the middle plate contracts, the hips connected to the upper plate cause the legs to pull inward toward the body while the hips connected to the lower plate cause the legs to swing outward away from the body. This motion is depicted in Fig. 4.2b More detail and an analytical formulation of the kinematics are given later in Section 4.2.2.

4.1.3 Prototype Testing - First Steps

To test the prototype, the degrees of freedom shown in Fig. 4.2 were actuated with 75 µm SMA wire. Actuator timing was controlled manually, and the prototype was run on ﬂat surfaces and an incline. Fig. 4.4 shows a sequence of frames from a video demonstrating the prototype walking in a quasi-static fashion on a ﬂat surface. Fig. 4.5 demonstrates the ability of the prototype to walk up inclines as large a 30◦.

44 1 2

3 4

5 6

7 8

Figure 4.4: Time sequence of steps from the robot prototype’s alternating tripod gait

45 Figure 4.5: The robot prototype climbing a 30◦ incline enabled by the addition of spherical rubber feet.

4.2 Design and Fabrication at Scale

Building on the success of the 2X scaled prototype, we proceeded with the design for the robot at the milli-scale.

4.2.1 Fabrication Process

RoACH uses a modified version of the SCM process that enables alignment of uncured parts to each other before curing. This slight modification eliminates one curing cycle (approximately four hours) from the process used to fabricate the MFI, and this represents a significant reduction in the total time required for a fabrication cycle. The SCM process for RoACH’s fabrication is depicted in Fig. 4.6 and the steps are outlined below.

1. The process begins with a sheet of uncured, pre-impregnated composite ﬁber laminate (S2- Glass).

2. Gaps are laser micro-machined into the composite ﬁber using a 25 W CO2 laser (Versalaser, Universal Laser Systems).

3. A polymer ﬁlm (PET) is placed on top of the laminate.

4. A laminate with the mirrored cut pattern is placed on top of the polymer layer.

5. The entire structure is cured.

6. The resulting ﬂat structure is released.

46 Figure 4.6: A step-by-step illustration of the SCM link and hinge fabrication process as implemented for the RoACH robot.

A palette of parts for the construction of RoACH is shown in Fig. 4.7. This palette demonstrates two distinct advantages of the SCM process. 1) It enables the fabrication of a robot with a large number of articulated joints for the same cost as a robot with a minimal number of joints. For example, RoACH’s design uses 57 joints. This is in contrast to robotic design with conventional materials and mechanisms in which the cost of each joint is significant. In addition, the ability to add an arbitrary number of joints to the design makes it possible to implement the kinematic redundancy strategy for SCM mechanisms suggested by the analysis in Chapter3. 2) Multiple copies of parts can be fabricated in parallel. The number of parts that can be made in parallel is limited only by the equipment used to produce them (ie. size of the laser cutting table, volume of the curing oven, etc.). These two factors combined make the fabrication of millirobots (from glass fiber composites, at least) inexpensive in terms of time and materials. From start to finish (including cure time and time required to populate the circuit board), RoACH requires approximately 16 hours to layup, cut out, assemble, and wire.

4.2.2 Mechanical Design

As discussed in Chapter1, size limitations severely affect the power and control budgets for mobile millirobot designs. At the milliscale, actuators can be expensive from a weight perspective (in the case of piezoelectric materials) and/or an efﬁciency perspective (the case with SMA). Thus, minimal actuation is a core guiding principle of robotic design at this scale. This minimal actuation constraint necessarily requires “programming” some of the robot’s gait into the structure via the kinematics. However, SCM has the advantage of being a near-monolithic fabrication process in which the cost of adding a single joint (and, hence, mechanism complexity) is essentially negligible.

47 Figure 4.7: A palette of ﬂat SCM parts showing 3 complete RoACH thoraxes ready for folding and assembly

48 This enables us to create very complex parallel kinematic designs for little to no added cost in manufacturing. (Notably, though, complex designs require more time to assemble by hand.)

Actuation

Given constraints imposed by the milliscale, we chose SMA wire to provide actuation for a number of reasons. Firstly, SMA integrates easily with the SCM process. Alignment and routing features can be directly machined into the composites to enable easy attachment and tensioning of the SMA wire. Secondly, SMA wire provides a force in a single direction. One difﬁculty of working with compliant mechanisms is their susceptibility to off-axis forces and undesirable moments as discussed in Chapter3. The single line force between points of attachment of SMA enables greater design freedom. Additionally, the target operating speed range (1 - 15 Hz) makes SMA viable. Lastly, it’s lightweight. One 80 mm actuator weighs less than 1 mg and can generate up to approximately 0.5 mW of mechanical power.

−2

−4 Gain (dB) −6

−8 −1 0 1 10 10 10 Log10 Frequency (Hz)

−3 x 10 1

0.8

0.6

0.4 Power (W) 0.2

0 −1 0 1 10 10 10 Log10 Frequency (Hz)

Figure 4.8: Frequency response (above) and output power (below) for a 25 µm (1 mil) diameter straight shape memory alloy wire actuator.

Dynamic characterization data for 25 µm and 37.5 µm SMA wire are shown in Figs. 4.8 and 4.9 respectively. The data shown were collected in a series of benchtop experiments in which the SMA wire was connected to a parallel beam ﬂexure spring with a stiffness of 170 N/m and actuated over a range of frequencies. Displacement was measured using an infrared emitter/collector sensor. The 25 µm wire resulted in a bandwidth (using the -3 dB criterion) of approximately 7.5 Hz with

49 5

−5 Gain (dB) −10

−15 0 10 Log10 Frequency (Hz)

700

600

500

400

300 Power (W)

200

100 0 10 Log10 Frequency (Hz) Figure 4.9: Frequency response (above) and output power (below) for a 37.5 µm (1.5 mil) diameter straight shape memory alloy wire actuator. peak power achieved at 6 Hz. The 37.5 µm wire demonstrated a bandwidth of approximately 3.5 Hz with peak power at the same frequency. However, because SMA is a tensile actuator only, it is necessary to provide a return force either with an antagonistic SMA or spring. With SCM, it is possible to design return springs into the hinges of the mechanism. Similar to tendons in animal muscles [25], this approach enables storage of elastic energy which is then returned when the actuator is switched off and the joint returns to its equilibrium state. The characteristics of the SMA actuators for our actuation conditions are summarized in Table 4.3. The SMA wire used in RoACH is 37.5 µm wire (Flexinol, Dynalloy Inc.), and it is attached to the robot by first crimping stainless steel hypodermic tubing to the ends of the wire and gluing the crimps to the robot’s body. Exposed portions of the crimps provide solder connections for wires providing the electrical power to heat the wire. The use of SMA is, however, not without its disadvantages. Lifetime of the wires can be significantly shortened by accidental overstraining or overheating. Care must be taken to ensure that preloading of the SMA does not exceed strain limits specified by the manufacturer [1]. In section 4.3, a drive scheme designed to heat the actuator quickly, but avoid overheating is described. This approach helps extend the life of the actuator. While manufacturer’s specifications claim lifetimes on the order of 106 cycles, a lifetime on the order of 104 - 105 cycles is a more reasonable estimate in this application.

50 Kinematics

RoACH’s alternating tripod gait is achieved through a combination of kinematics and basic feedforward, clock-based control. A schematic depiction of the gait and a map of the actuator inputs corresponding to the gait positions are shown in Fig. 4.1 and Table 4.1 respectively.

30mm

20mm

Figure 4.10: A CAD model of the robot showing the body-ﬁxed basis used for kinematic analysis of the motion of the limbs. The top plate is shown nearly transparent so the inside of the robot may be seen.

RoACH’s kinematics can be reduced to two translational degrees of freedom which are appropriately coupled through a central 2 DOF translating element to fourbar linkages that map those translational inputs into angular outputs. The design is ground-free in the sense that any rigid link could be used to establish the body coordinate system. RoACH’s kinematics can be understood by examining two structures and their interactions: the central plate (shown in blue in the CAD model in Fig. 4.10) and the two outer plates of the body (shown in cyan, with the top plate transparent in Fig. 4.10). The outer plates are constrained to move together in the horizontal plane, but are free to translate vertically with respect to each other. The middle plate is designed to contract laterally in the horizontal plane and is coupled to the upper and lower plates through fourbar linkages at the “hips” of the robot. From Fig. 4.10, when the central plate is moved forward along the anteroposterior axis of the body with respect to the outer plates, the magenta tripod swings back and the yellow tripod forward. When the central plate is contracted along the mediolateral axis of the body the magenta legs are lifted, and the yellow legs are lowered. These motions are also depicted individu- ally in Fig. 4.11 while Fig. 4.13 shows the conﬁguration of the SMA wire actuators used to control the two degrees of freedom. The two links of each hip fourbar that remain parallel to the ground are hinged one to the middle plate and the other to either the upper or lower plate depending on which tripod it belongs

51 (a) First degree of freedom - abduc- (b) Second degree of freedom - ﬂex- tion/adduction of the legs (rear view): ion/extension of the leg (side view): model and actual robot (with only one tripod model and actual robot (one hip only) attached)

Figure 4.11: Crawler kinematics depicting two independent degrees of freedom to. Hips which are hinged to the top plate will cause their respective legs to be pulled toward the ground when the middle plate contracts laterally and those hinged to the bottom plate will be lifted from the ground. Likewise those hips hinged to the top plate will swing the legs forward when then middle plate moves forward with respect to the top and bottom plates while the hips hinged to the lower plate will swing the legs backward. The kinematics can be written explicitly using the body ﬁxed basis shown in Fig. 4.10 in which the origin of the basis is located at the idealized point on the top plate where the middle leg attaches. The x1 input represents the contraction of the ﬁrst SMA in the e1 direction and swings the legs forward and back. The x2 input represents the contraction of the second SMA in the e2 direction and is responsible for swinging the legs inward toward (or outward away from) the body to engage or disengage the tripods with the ground. The position of the end of the middle leg with the coordinate basis attached at its hip can be computed from a simple rotation matrix:

cos θ cos ψ sin θ sin ψ sin θ e1 r = L sin θ cos ψ cos θ cos θ sin ψ e2 (4.1) leg −    0 sin ψ cos ψ e3 −     1 x2 ψ = ψ + sin− ( ) (4.2) 0 2l

1 x1 θ = θ + sin− ( ) (4.3) 0 2l In Eqns. 4.2 and 4.3 l is the length of the vertical link of the hip fourbar as pictured in Fig. 4.16

52 and ψ0 and θ0 are the unactuated angles of the hip. Once the position of a single leg is known, the positions of the others are trivially determined due to the parallel nature of the design. Kinematics were also simulated using Matlab and the resulting visualization of the foot paths in the body-ﬁxed reference frame is shown in Fig. 4.12.

Figure 4.12: Output from a simulation of the robot kinematics showing the closed paths along which the feet travel with respect to the body.

Flexure linkage design

Each degree of freedom in the robot begins its actuation cycle at one extreme of its output motion and must be returned to this position at the end of one actuation cycle by the flexure return springs. Because our actuators provide primarily line forces aligned with the most compliant directions of our linkages, we have chosen to use the PRBM to get approximate linkage stiffness for the purpose of sizing our flexure hinges and our actuators. Using the short flexure approximation [47], recall that the expression for the angular spring constant of a simple flexure hinge is:

Ebh3 K = (4.4) θ L In Eqn. 4.4 E is the flexural modulus, b is the width of the flexure, h is the thickness, and L is the length. The first return spring works in opposition to contraction of the middle plate shown at the top of Fig. 4.13. It consists of a compliant approximation of a Sarrus linkage - a spatial linkage which provides a single translational degree of freedom. The second degree of freedom is provided by the parallel fourbar linkage at the hip joint shown in Fig 4.16. The equivalent spring “constant” (the

53 Sarrus Linkage/Return Spring

SMA Route

3mm

SMA Route

Figure 4.13: Above: The central contracting plate showing the route for the actuator responsible for abduction and adduction of the legs. Below: The bottom plate depicting the routing of the actuator connecting the bottom and middle plates and used to ﬂexion and extension of the legs.

54 spring force is a nonlinear function of linkage displacement) for these linkages can be determined using Castigliano’s second theorem as follows:

∂U ∂s ∂s ∂x = F = F (4.5) ∂q1 ∂q1 ∂x ∂q1

In Eqn. 4.5 U is the potential or strain energy stored in the linkage, q1 is a generalized coordinate (angular displacement of a link in our case), s is a generalized output displacement, and x is a linear output displacement. In our case, s and x are equivalent, so we solve for F as a function of x and differentiate to give the expression for the equivalent linkage spring constant, K:

∂2U K = (4.6) ∂x2 In the fourbar linkage it is also possible to induce the linkage to rest in a biased position such that the joint angles are not all 90◦. This is accomplished by stiffening the ﬂexure hinges at opposite corners. For a parallel fourbar with 2 joints catercorner from each other having a stiffness, K1, and the other two having stiffness K2 the rest angle, θ is given as:

π θ = (4.7) 1 + K2 K1 Thus it is possible to affect an arbitrary bias angle and therefore set the return point of the leg for the beginning of its stance phase simply by adjusting the stiffness of a ﬂexure (either through material properties or geometric changes). For a fourbar with an arbitrary bias angle induced by a stiffness differential in two of the four joints and loaded as in Fig. 4.14, the expression for the spring constant resulting from Eqn. 4.6 is:

n + n K = 1 2 (4.8) fourbar m

(x + d) 2 n = 2El 1 (I + I ) 1 2 1 2 (4.9) − r − l

1 x πI2 n = 2E (x + d) cos− + cos (I + I ) πI (4.10) 2 l I + I 1 2 − 2 1 2

(x + d)2 m = lL (d l + x)(d + l + x) 1 (4.11) − s − l2

πI d = l 2 (4.12) I1 + I2

In Eqns. 4.8- 4.12, I1 and I2 are the area moments of inertia for the ﬂexures represented as K1 and K2 respectively in Fig. 4.14, E is the modulus of the ﬂexure material, and L is the length of

55 all the flexures. In practice, the stiffness of each flexure joint could also be adjusted by changing L. Fig. 4.15 shows the force displacement curve for a biased fourbar with the same flexure and link length dimensions used on the actual robot.

K (E,I ,L) 1 k11 Kk2(E,I2 2,L) FF θ2 l l θ1

x K1(E,I1,L) K2k(E,I2 2,L) k1 Figure 4.14: A kinematic model of a fourbar linkage using the pseudo-rigid body model to approximate joint stiffness.

0.06

0.05

0.04

0.03 Force (N) 0.02

0.01

0 0 0.5 1 1.5 2 2.5 3 !3 Displacement (m) x 10 Figure 4.15: Force displacement curve for a biased fourbar linkage loaded only in the x direction.

Fig. 4.17 shows half of a Sarrus linkage where the linear constraint imposed by the non-planar portion of the linkage is instead replaced by a kinematically equivalent sliding joint. For the Sarrus linkage conﬁguration shown, the stiffness is:

x 1 x 3EI 3EI 1 2l cos− 1 2l KSarrus = − − (4.13) x 2 3/2 2 − 2 x 2 2l L 1 1 2l 2l L 1 1 − − − − 2l

56 K2 K1 l F

K1 K2

3mm Figure 4.16: Above: A model of a biased parallel fourbar (shown relaxed and displaced) using ﬂexure hinges as springs with different rotational spring constants. Below: an actual hip fourbar in equilibrium position on the robot showing the biased return spring for forward and backward swinging of the legs.

Fig. 4.18 shows a force displacement curve for a Sarrus linkage of the same dimensions as the one used on the RoACH robot. We note that when α = 0, the mechanism is singular, so a stiffness cannot be computed at the position. In practice, we always operate the Sarrus linkage with some initial displacement to prevent it reaching a singular conﬁguration.

Kθ(E,I,Lkr ) l l F α Kkθ(rE,I,L) Kkθr(E,I,L) x Figure 4.17: A kinematic model of half of a Sarrus linkage using the pseudo-rigid body model to approximate joint stiffness.

SMA Kinematic Constraints

It is also important to note that the displacement limitations of the SMA wire must be considered in the kinematic design. For example, to achieve our desired angular displacement of the legs, we require approximately 3 mm of linear displacement from each actuator. By conservative

57 !4 x 10 4.55

4.5

4.45

4.4

4.35

Force (N) 4.3

4.25

4.2

4.15 0 0.5 1 1.5 2 2.5 3 !3 Displacement (m) x 10 Figure 4.18: Force displacement curve for a Sarrus linkage loaded in the x direction.

estimation, SMA wire is capable of 3% strain, implying a requirement of 100 mm length wire. In theory it would be possible to achieve the same displacement output from several SMA wires connected mechanically in series, but electrically in parallel and avoid issues associated with the high electrical resistance that accompanies long wires. However, for simplicity, we have chosen a single continuous wire for each of the actuators. The electronics design ramiﬁcations of this decision are discussed in more detail in the following section.

4.3 Electronics Design

For very small robots, integration of power and control electronics is a crucial design consideration. Since the selection of lightweight batteries at this scale is limited, the battery represents a high percentage of the overall robot weight. Likewise, since custom integrated circuits are quite expensive, lightweight PC board construction is the practical solution for electronics design. Nonetheless, the electronics are still relatively heavy, representing a signiﬁcant portion of the robot’s weight. Careful selection and design of electronic components can decrease overall robot weight signiﬁcantly, improving performance.

4.3.1 Control Board

A lightweight circuit board is needed for controlling the SMA actuators as well as for steering, communication, and potential future sensing needs. A custom 402 mg control electronics board (shown in Fig. 4.19) was manufactured using, 25 µm core FR4 ﬁberglass. The board utilizes surface mount components and features a connector not only to the power electronics but also to a

58 System Component Mass (mg) Blank Board 180 Control PIC Processor + LEDs 97 Electronics Regulator + Battery Monitor 22 IrDA 103 Blank Board 124 SMA Power DC-DC Converter 153 Electronics Drive Transistors 10 Power Supply Battery 847 Structure Skeleton + Joints 700 Actuators SMA Wires + Crimps 2 Wiring Wiring, Solder, Trim 160 Total Mass: 2398 Table 4.2: Mass breakdown for the entire robot

sensing “daughter” card. None is presented here but several daughter cards exist for such sensors as accelerometers, gyros, or a compass. This control board is a more recent revision of the control board used in [92], which demonstrated autonomous glider navigation toward a light using an infrared sensor.

4.3.2 Power and Processing

The control board contains a 10 Mips PIC LF2520 processor running at 3.3 V through a ZXCL330 low dropout regulator. The PIC processor was chosen for its internal oscillator to save weight and for its low power consumption. Red and green LEDs indicate battery voltage status using a MN1382 battery monitor. RoACH is powered by a 20 mAh Lithium polymer battery (Model# 301218HS20C, FULLRIVER Battery New Technology Co.) that weighs 850 mg (after trimming).

4.3.3 Wireless Communication

The board contains a complete implementation of the bidirectional IrDA infrared communication standard common in PDAs and laptops. Infrared communication uses very little power compared with RF technologies like Bluetooth. The lack of antenna also decreases the weight of the module at the expense of range. Communication uses the PIC’s hardware UART and communicates with a PC at 115.2kbps via a serial port and IrDA converter.

59 Power Con- I. ELECTRONIC DESIGN 14mm IrDA nector For very small robots, integration of power and control electronics becomes an important design criteria. Since the selection of lightweight batteries at this scale is limited, Power the battery will be a very high percentage of the overall Processor robot weight. Likewise, since custom ICs are quite expensive, Cond. lightweight PC board construction is the most practical solution for electronics design for prototyping but is still 24mm relatively heavy and will be a significant portion of the Sensor robot weight. Therefore, any gains that can be found in the design of these components can decrease overall robot weight significantly and hence improve performance. Top Bottom A. Control Board Fig. 1. Top and bottom of the control electronics PC board. A lightweight control board is needed for not only dFigureriving 4.19: Top and bottom of the control electronics PC board. the SMA actuators but for other uses such as steering, communication, and sensing behaviors. A custom 440mg Wire Diameter Force Out- Resistance per Maximum Pulse (mil) put (g) Length (Ω/in) Current (mA) control electronics board (shown in Fig. 1) was manufac- 1.0 7 45 60 tured using lightweight, 1 mil core FR4 fiberglass. The 1.5 17 21 100 board utilizes lighweight surface4.3.4mount co SMAmponent Powers and Supply TABLE I features a connector not only to the power electronics but TWO FLEXINOL SIZES AND ASSOCIATED CHARACTERISTICS. MAXIMUM also to a sensing “daughter” card; none will be presented SMA wire is a thermalP actuator,ULSE CURREN actuatedT ESTIMATES byARE resistiveFROM OUR O heatingWN EXPERI withMENTS the. power requirement for here but several daughter cards eeachxist f actuatoror such se givennsors a by:s accelerometers, gyros, or a compass. This control board is 2 the next generation of the control board used in [1] that P = I R (4.14) demonstrated autonomous navigatwhereion tow Rard isa theligh resistancet using th ofe n theomin Flexinolal battery (SMA)voltage wire.of 3.7V The. A SMAsafety f poweractor w supplyas should be capable an infrared sensor. included to make the target output voltage of the SMA power 1) Power and Processing: Theofco drivingntrol boar ad varietycontains ofa wireconv sizes.erter 13 Available.6V. This is Flexinolthe highe wirest volt diametersage that is n andeede thed associated resistance 10Mips PIC LF2520 processor rupernnin unitg at length3.3V thr (fromough a www.dynalloy.com)to drive either of the arewire showndiamete inrs Tablein Tab l4.3e I.. Maximum ZXCL330 low dropout regulator. This particular PIC proces- output power of the proposed DC to DC converter is 0.83W sor was chosen for its internal oscillaWiretor to Diametersave weight (µanm)d fForceor a sin Outputgle chann (g)el of tResistancehe 1.5mil wire (;Ωf/in)or the twMaximumo channels Pulse Current (mA) its low power use. The board also has a red and25green LED on the robo7t it is 1.66W. This targ45et power output is an upper 60 to indicate battery voltage status using a MN1382 battery design bound only; in reality the robot does not have 200mm monitor. 37.5 of Flexino17l and both actuators wo21uld not be pulsing to their 100 2) Wireless Communication: The board contains a com- maximum drive current at the same time. plete implementation of the bidirecTabletional I 4.3:rDA in Twofrared Flexinolcom- sizesA stan anddard associatedboost conver characteristics.ter topology was c Maximumhosen as the pulse current estimates munication standard common in PareDAs fromand la ourptops own. Infr experiments.ared DC to DC converter. A Coilcraft LPS4018 15.3µH surface communication was chosen for its very low power use mount inductor that weighs only 97mg was used for its compared with RF technologies such Theas bl maximumuetooth as w lengthell lig ofht SMAweight wireand hi ingh thecurr designent toler isanc approximatelye (0.86A, 10% dr 100op) mm. For 25 µm wire, as its low weight (lack of an antenna). The communication for the high powers needed here. Switching is realized utilizes the PIC’s hardware UARTtheand resistancecommunica istes 177to aΩ. Tovia drivean LT the3580 maximumboost conv oferte 60r I mAC ca pulsedpable of currentswitchin throughg this resistance, a PC at 115.2kbps via a serial port voltageand IrDA ofco 10.6nverte Vr. is needed,frequ meaningencies from a DC200k toHz DCto converter2.5MHz; a ishig necessaryh speed, hi togh boost from the nominal power schottky diode was used for switching (UPS120). B. SMA Power Supply battery voltage of 3.7 V. A safety factor was included to make the target output voltage of the SMA power converter 13.6 V.R Thisesisto isrs thegove highestrn both th voltagee switchi necessaryng frequency toan drived the ou eithertput of the wire diameters SMA wire is a thermal actuator; it is simply a resistive voltage, making the circuit easily customizable for different load and can be actuated by passingincu Tablerrent t h4.3roug.h Maximumthe wire. S outputMA wir powere diamet ofers. theAn proposedefficiency pl DCot for to1M DCHz a converternd 2MHz is 0.83 W for a single Therefore, the power requirement fchannelor each ac drivingtuator is s theimp 37ly µmswi wiretching orfre 1.66quenc Wies a fort 13 both.6V ou channels.tput voltage Thisand 3 target.5V inpu powert output is an upper 2 given by P = I R, where R is thedesignresistanc bounde of the only;Flexin bothol v actuatorsoltage is sh areown notin F pulsedig. 2. to their maximum drive current at the same time. (SMA) wire. The SMA power supply to be designed should The boost converter was implemented on a 1 mil core FR4 be able to drive a variety of wire sizeAs. A standardvailable Fl boostexinol convertercustom PC topologyboard (like wasthe c chosenontrol bo asard) the. It w DCas fa tobri DCcated converter. A Coilcraft wire diameters and the associatedLPS4018resistance p 15.3er uniµt lHeng surfaceth to mountmatch th inductore power co weighingnnector on 97the mgcon wastrol b usedoard a fornd i itss light weight and high (from www.dynalloy.com) are shown in Table I. shown in Fig. 3. Its total weight is 340mg. The maximum length of SMAcurrentwire to tolerancebe used in (0.86the A, 10% drop). Switching is realized via an LT3580 boost converter IC ca- design is approximately 100mm. pableIf this ofis 1 switchingmil wire, frequenciesthe C. SMA fromDrive 200Meth kHzod to 2.5 MHz. A high speed, Schottky diode was used resistance is 177 Ω. To drive the mforaxim switchingum of 60mA (UPS120).pulsed ResistorsThe boost co governnverter w bothas use thed as switchingthe biasing v frequencyoltage for an and the output voltage, current through this resistance, a voltage of 10.6V is needed, NMOS transistor (FDG6301N). An optional current sensing meaning a DC to DC converter is necessary to boost from resistor is routed on the PC board to an A/D input of the PIC 60 making the circuit easily customizable for different SMA wire diameters. An efficiency plot for 1 MHz and 2 MHz switching frequencies at 13.6 V output voltage and 3.5 V input voltage is shown in Fig. 4.20. The use of the DC to DC converter could have been avoided by connecting several SMA wires mechanically in series, but electrically in parallel. The 3.7 V battery is capable of sourcing the required 60 mA of current (for a 25 µm SMA) into a 61 Ω resistive load. With a resistance of 1.8 Ω per mm, a 100 mm actuator would need to be broken into at least three segments to be driven by the battery with no boost circuitry. The series mechanical/parallel electrical arrangement adds significant complexity to the design motivating our design decision to use the DC to DC converter. +Vbias 85

80 +SMA )

% 75 ( -SMA 1 MHz Switching y

c 2 MHz Switching n 70 e i c

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Figure 4.20:F Miniatureig. 2. Min poweriature p supplyower su efﬁciencypply efﬁcie forncy 1f MHzor 1M andHz a 2nd MHz2MH switchingz switching frequencies. frequencies.

Fig. 4. Transistor drive scheme for SMA. The boost converter was implemented on a 25 µm core FR4 custom PC board (like the control board). It was fabricated to match the power connector on the control board and is shown in Fig. shown in Fig. 3. Its total weight is 340mg. 4.21. Its total weight is 340 mg. 70 Spike - 1 Power Connector 60 ) Spike - 2

4.3.5 SMA Drive Method A

m 50 (

The boost converter was used as the biasing voltage for an NMOS transistor (FDG6301N).t An n

e 40 optional current sensing resistorLT358 is0 routed on the PC board to an A/D input of the PIC processorr Inductor r for future implementation of closed loop current control. The gate of the NMOS was drivenu by

C 30 Normal - 2

the PIC with a PWM signal (19 kHz) to govern the current through the SMA wire. A “spike”t

18mm u plus “normal” current drive method was used to increase efﬁciency of the actuator. This approachp t 20 achieves initial actuation with a large spike in the current drive that quickly heats the wire.u This Schottky Diode O 10 Normal - 1 61 14mm 0 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 Time (s) Fig. 3. Miniature boost converter pc board. Fig. 5. Typical current drive proﬁle showing spike and normal hold current (1Hz).

C. SMA Drive Method D. Communication/Software The boost converter was used as the biasing voltage for As mentioned earlier, the control board utilizes IrDA for an NMOS transistor (FDG6301N) as shown in Fig. 4. An wireless communication. Both the gait control PC and the optional current sensing resistor is routed on the PC board robot control board implement a cyclic redundancy check to to an A/D input of the PIC processor for closed loop current detect communcation errors. A GUI was written in Matlab control (not implemented in this work). to control robot gaits; the software is able to communicate The gate of the NMOS was driven by the PIC with a PWM with the robot at all times and can change gaits (if needed) signal (19kHz) to govern the current through the Flexinol each step cycle. However, the robot is driven in an open wire. A “spike” plus “normal” current drive method was used loop, teleoperated manner for the work discussed here. The to increase efficiency of the SMA actuator. In this method, step profile is only changed every several steps for turning initial actuation is achieved by a large spike in the current or stopping. drive that quickly heats the wire but would damage it if it REFERENCES were held on constantly. This spike is followed by a lower [1] R. Wood, S. Avadhanula, E. Steltz, M. Seeman, J. Entwistle, current level (normal level) that keeps the wire heated but A. Bachrach, G. Barrows, S. Sanders, and R. Fearing, “Enabling would not normally heat the wire very quickly if it were technologies and subsystem integration for an autonomous palm-sized used as the main drive. glider,” IEEE Robotics and Automation Magaznie, vol. 14, pp. 82–91, 2007. Both actuators were heated in this way. To achieve steering, the two channels were driven with a variable phase difference between them. Current profiles for the two actuators for a typical drive profile are shown in Fig. 5. s f F f f i u w c w d i t w s c t o a C r n n o o o e h i u o n i s r p e g . g o e i r g i i r q o e r n t T T B r t . a v e i r n u r , i u d e i w N S n t

e Efficiency (%) a 2 n n o e e h h a o a e l r . t M . c d n c n o n l 6 7 8 5 6 7 8 n M e e l h t a A t A h r t c h e t l a 0 0 0 5 5 5 5 y s e h e a A i g e ( M i 0 n b O l e / p a 1 c ( a n “ l b l a a D t s t o o n e d i i s t t h 9 c w s . S D c t e n c u v t o e o e F p u e k t t i L a q i o a a o r e u s w t i r i n H t n o l t u e i T t k t l g F n m r a r u v p i i e o f c f e a i e i d z . o t r m 3 e c fi g ( c u e h e o r a n n 1 ” t c ) r n n 5 . o 3 m k h n t c a i r p i t s o 8 p M o 8 n 3 v p n . s l e i n t i o l o n m i a y e . t s o r s 0 v e l e w s n I e h f m l w u N s d n e t m t e l m o e t e t e s h n p s r c y t g a r h M a e M r a l r h i m e t y e s r s c o n v “ o r t l e s o i e i n a o e h h u n n v ( e O t d . n O 0 r w fi o t t l l F p t i i e . e o P s g P e y . e a C a e h S f p u l 5 w a D e r t r . d v I e l v o l u e n u t t m t r C y r t e a w p r w e h G a T e w h r e e a i l s d e t u t a e r n t ) s e a h r e 6 h p e f e e d 1 b h l t i e fi s r i u d ” 3 e r o n i s b r w 4 t e t s S c g P o h h i o t s 0 d t s C y i m c 2 v 1 o M i h i e e s c o w a i s h 1 r c n r u p t e n s r o d e t i M p t w M m e o u N a v c r n i r c A s n i r y r t w w i i o r o k H e H e s k a ) h s s e n e b r l n w fi n o r e f e a s z e o n z n i e u v a o 3 a e s r r l r n v r e i c t ( c r b t S e S o g s p t t i 4 k i r W) 1 t e t f t h w h s w d s w y 1 t n u e o s w u 0 o ) e . r M e s r i 0 t . i t r y r a f a S m r f t t a t i h e h F o s c t c h o v o H t y p c h d b h r c p o h u o i g e r I c v e . l q o h z g i e i l i w l i n l r n . n a o a u o k t T u d b . o o P . m a g d g h r s n s o w i w e n g t o n I i I 5 u c i e e t a pk sfloe yalwrcretlvl(omllvl htkestewr etd nillustration An . 4.23 heated. wire Fig. the in keeps shown that is level) scheme (normal drive level the current of lower a by followed is spike e d d a C n h n k . e k r i c d t i i b a a t d r g t 2 y h n d n s f F f f i u w c w d i t w s c t o a C t l h e w c . w m l t o y M t r n n o o l o o e e h h i e h u o n i h o s r p e D r v b t o F g . g i d o a e i i H h r g i e h i r i o t q o e p i o s r n y t B T T r e P g i t . h a v i e e f z i r n e u p w a g r , i h u o l v d e i w N e S a C n t F m t e Efficiency (%) a 2 n n o c e s e h h a . a a a d o a e e a i l r 1 a t . c c w l t t t M . c e s d e i g n e c n o n l e s u 6 7 8 5 6 7 8 n M . 4 e e l u u P h t a b l t i e s d A 5 t A h x t r e t t o a w c h . r r e u WM t t l a c o 0 0 0 h 5 5 5 5 y s e h i e i r r a e A t w i g d h e ( M s f n a i b o 0 n o e e A e b O l f e / e p i a i e 1 c ( a n “ r o l b l n o u n n r e a a d D r t f s r t i o o d n d n e e d g s i i s r r - - t t h l t t t t , 9 c w s . 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e Efficiency (%) a 2 n n o e e h h a o a e l r . t M . c d n c n o n l 6 7 8 5 6 7 8 n M e e l h t a A t A h r t c h e t l a 0 0 0 5 5 5 5 y s e h e a A i g e ( M i 0 n b O l e / p a 1 c ( a n “ l b l a a D t s t o o n e d i i s t t h 9 c w s . 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S m , 0 o o e E e y l c l i i o i s F p b n n r u s E h c c c m S t b a i i u , o p e a l n a g . k 0 t e o o R r n a u d t . u c t e r G e . . d o a n i t e n a 4 1 t A d o g b . n r l e . t i s H - a v d y n o a t i c u a m B a l o t 0 i d / a b 1 o r l i d t S i T a c n r c . l s t s a h w m i 2 r s i t r o y i h v S ; r a n e i a o s o e e f e a m n a p n p r t n t n w t R 0 v n r e u , n s p w l h i e r . m e i d l e . s e k r g s a E t e o 3 r , o a s m r t h N , e B e A o r fi F , i r T f e S n s s d 63 r u a l e e E o 0 o o - E e . o t . t t i ◦ n n d e r . o t R h m r c . hspaedfeec ewe h w actuator two the between difference phase This . f e s 4 g 2 h r m d A S m t o t e h N i E v r t v w a e S o a a a n h a e N e c n 0 t o t w t t r G h t i a e l r d i e a ( . o c C r r o s o y i l 5 s r e e n c m n o n - t y U n b e w r E ) z h g s l c , s o f 1 e g S a + M I , c 0 o o l i s e me b t l h a i s . p r r a M v a w c V 6 o i i a k n g - i a k e t a b a . n s a d a n r e f r i b 2 0 r d o e z s c g a a + - a d n a S l r . d d R i R o l e s S n e 7 u i s e w S S r e u n . d t e c M u s s i , o g M M n t t ma v r u t en 0 n t F n r o v a i e d i e A o s A o o e t . o l A i p 8 t a n r n s a s i mo l t l c . e ma . z s r s e , n i o n e 1 P d n 0 c i u ( f m s 4 J l n g y C . i o s . i , , 9 h h f n m r I p o p c e a r “ E n a l a p h D r n u E t d M l n . e n u e m- e n n t c A e 8 . d w r c a o a u i 2 d n b s c k T i r t p – i s e i l r t l a f z i t n 9 e e a h h d o n l t e t n 1 e n b g o e e e d g ) r t , , f F l e w t d r w D i a w u b l i c m w F w i p s r o o n o t n o t a u i u e s r e l g e a a i i . i w o e v o b e c q e i t r t A t r t . t p S y s c f e t a x r p h h u h o c d e h C e

h Efficiency (%) 2 w . e o c e e r s i e l , t i u i r . o p a i n n n e 6 7 8 5 6 7 8 t n c t o s T t s a o s a s t h c c m d s r o t c t 0 0 0 5 5 5 5 s t m e o s P t b e u c i w o w i o M s r e 0 e o , u e l l o d d o o WM l e l fi m n s e p t r a r e e o d i w i n . r c n r l m h n o r t n l c r i t o r u r o e e t i f e v L e o i p p n h k i r c a o I i m b o n r n m t o o e t I T h e y o i ) F e i l u i o r i m t h . i e . a b n r u s s a c i c t . r m n 3 t b g v t a e l e i l o e s 1 l c T n a h a A c 5 n g . e t e l e o o e a u d l t 8 c p t a e i d a r h v 8 o 3 n . o a n i n t o n e a c m d t s . o e e g “ 0 d r a s r w l e a t t i o t H a m l d y n s e a l i i q u c e d a m s m e g l n f o p d i / M ( a ( e o u a r l e r r f a t S a i n c i n 1 l i e t r s a t w m i l s k c t i v i t o i h i i l c o e 9 ; l r u n n e e i O o 0 o n o o e y e e f e e m k a f p p i r k r n P t n . a t ” n g o fi n . v r n m , o p n u l 5 w l h H t e h r . e y o e , p s l u e c c f g t B p t e o y r i a a e s t p r m w z r h i e t B s e a l e r t , e l a h h r o c e u ) u f e h s s d n a e e e n f t o 1 b e o b e u a t l t . o s t fi t e n r t h n e c o a o t h 4 c e r c r t l c f e P h d t v A h o “ t y r y t h t o C t e N i m 2 w v 1 s g e t e s w a w e n o . e a i n h t h esaedie otnosya H,afeunyjs eet h d on o 37.5 for point 3dB when the minutes beneath 9 just over frequency of a time 3Hz, running at wire. or maximum continuously length a driven body indicate are 1 tests legs approximately life the at battery crawling Experimental robot the cm/s. show 3 results best Our experimentally. tuned e n c n o t c o o M l e M w M m o o r G T t e h t c a e ) r c w a t v t n v f c h r o r y y o e u r H e H e n h o c O m e y U t n e e b e w i n c n t a h r f z h i z d r l r c h e v o s o o s S g t I c a e a n o w a e l i c ( e o e r S t S b t n h a l i t s r b W) 1 h n r t v s w w ” c r w 1 v t t w o i t i o a k e S y . k p h e s n i r e r t M e a i 0 i a r n s a r S c e o r a t m t e i e M r e r b c c r d e k s c g w H u a p a s l e c y l h d h l c d d b i I s o c e l e r p h z A e i i s e ( w n ( d u l n r n r u u n q s b o c m n a u w s r a e i g d i g r g r t n t t u v o r e u n s t r t o i t n a r r o u t a i e t e v a g o d i d e i e w s o h c t t k t l r i c c p f r e h e r t a n n s i d t t l e 2 c y o e e k d e t z o s i s n u e n e . t o M o m n l s e l r u P d a w a c i d l t y D r o i ( f m p b s n o H h t l t y v C p i o r n i e d i o i y h f w i r i e n i z r m l k r I d o v r o f c e e a e r . e s n e s d a t e d m h u D m r n 1 u t M h i w t d e n i a u h I e n e t e e g . n i n e n A i e e m e 5 . d r e n e c t h o a w i b t c w n d n a c r k P T h t iue42:Te24gRAHsadn vraU quarter US a over standing RoACH g 2.4 The 4.25: Figure t t p h a y i e i t t t t t l e a h h f i i o n n I g e e e h h h a h d o n t r t t i i C d n d d g g b o h e e e e e e e a g s s ) r o f d t a b e l t a t m p c [ 1 h e h h o n c o r e e l ] g m e e a a i r C P t p g n T d B u w L s s n t m o R 2 g t A a e t e o o o d t s s e a r a i 0 l o a g n w iue42:Eetoisbokdiagram. block Electronics 4.24: Figure t w . i o c . m n o o l i P t t 0 h a d d ( r e p h i d i t n d v l e o t 7 d n o e i B c e e n e W e l p i r n . r . i r e l l h . g n n o r s a i n , n i t y ” o e g o y p d l c n i A g , h t o g g o t e h w w o I g e g h d fi i E r s v s t s d o i s a . i h i , e i i e E e n S n t c l s m g n w r e s A h l a E h s i e w S t r T s f , p t w a u p a e o w c o . t h o h t R n r i p s h r n l r t u G F c t i I i d e u e o y n h u i l c t A e p r n i c p b . s u i c r i g i i a D h s I o v e o u n d t o . n c o h . g s u t o I h B a i d t d p r A h o e g b d n 4 T i i r d . F u r ; a c n s p s o d r . , h g o t r s h t i y T i s g r o e o a T u . d e C d g m u s o i i a n r i ) s r n t s d i E e . 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Table 4.4: Summary of the physical parameters for RoACH

Onboard Speed (body Robot Size Mass Power/Control lengths/sec) Silicon µbot [28] 1.5 cm 83 mg No 0.4 Mesoscale Quadruped [38] 9 cm 104 g Yes 3 LIPCA Hexapod [99] 12 cm 35 g No 1.5 HexBug [49] 5 cm 15.6 g Yes 1.2 ARRIpede [68] 1 cm 3.8 g Yes 0.15 RoACH 3 cm 2.4 g Yes 1

Table 4.5: A comparison of similarly sized legged robots

4.4.1 Turning Control

To achieve turning, differential leg lengths were used. Both middle legs are slightly shorter than the four corner legs. During front/back swing actuation, the lateral degree of freedom is actuated before the swing ﬁnishes. This momentarily places the four corner legs on the ground simultaneously without the middle legs engaging with the ground. Since the legs on the left side are moving in the opposite direction to the legs on the right, a moment is applied to the robot. Slipping on one or both of the sides will occur, causing the robot to turn. As seen in Fig. 4.26, both left and right turning can be controlled simply through timing the actuation of the lateral (tripod switching) degree of freedom. To our surprise, gait parameter exploration also revealed that with 1 zero forward velocity, the robot is capable of turning at a rate of approximately 1.5 rad s− .

4.4.2 Energetics

One major focus within the study of locomotion in organisms is energetics. One common approach to assessing the efﬁciency of an animal’s mode of locomotion is called metabolic cost of transport [7]. The metabolic cost of transport is the measure of total energy required by an

65 Figure 4.26: Composited frame captures from video demonstrating controlled right and left turning

organism to move one kilogram of its mass one meter. The metabolic cost of transport can also be written:

Power Metabolic Cost of Transport = (4.15) Speed Mass ∗ This metric can also be applied to engineered systems like the RoACH robot. At maximum speed, RoACH consumes approximately 1.3 W of power. At the maximum walking speed, this implies a metabolic cost of transport of nearly 18,000 J/Kg-m! Fig. 4.27 puts this number in the context of biological systems. Though organisms necessarily become less efficient as their size decreases, RoACH’s metabolic cost of transport is still more than two orders of magnitude larger than similarly sized biological systems. The primary cause of RoACH’s high cost of transport is the low efficiency of the shape memory alloy actuators. Previous studies [85], [63] have reported theoretical efficiency limits in the range of 3 - 13% for a single straight wire and ideal Carnot efficiencies of 8 - 15% for the same configuration depending on how output work is calculated. However, experimental reports of 1% efficiency (or less) are common in the literature [50]. This severe performance limitation will motivate the approach taken in the following chapter of using a single DC motor as a main power drive and shape memory alloys as lightweight steering actuators.

4.5 Concluding Remarks

In this chapter, we have presented mechanical and electronic designs as well as experimental results for an autonomous 2.4g hexapedal millirobot called RoACH. The robot is enabled by the SCM process for millirobot fabrication. We have demonstrated straight walking and controlled turning at speeds up to 3 cm/s. To the author’s knowledge, as of 2010, RoACH was the smallest and lightest autonomous legged robot. Though RoACH represents a signiﬁcant step forward in the design and fabrication of legged

66 RoACH

Figure 4.27: Data from [33] showing typical costs of transport for organisms over a range of sizes. Metabolic cost of transport for RoACH is shown in red and is approximately 18,000 - more than two orders of magnitude larger than similarly sized legged organisms. mobile millirobots, efficiency of locomotion at this scale continues to be an issue. While shape memory alloy actuators are a good choice for actuation from the perspectives of ease of integration and control as well as power density, they are remarkably inefficient, especially compared to standard electromechanical actuators like DC motors. The next chapter introduces a hexapedal legged robot design approximately one order of magnitude larger (in mass) than RoACH, that takes advantage of dynamic locomotion and more efficient actuation and kinematics to be several orders of magnitude more efficient.

67 Chapter 5

DynaRoACH: Design and Dynamic Maneuverability of a Hexapedal Runner

In this chapter, we present the kinematic design of a lightweight and efﬁcient SCM-based transmission for driving a small legged robot from a single DC motor. Compared to the RoACH robot of the previous chapter, actuation using a DC motor enables far more dynamic and efﬁcient locomotion than with SMA actuators. The single motor design also reduces the mass of the robot compared to one motor per leg designs like RHex [75] and yields a feedforward clock-driven gait. Coupling the transmission with passive compliant legs which exhibit viscoelastic damping results in stable dynamic straight-ahead running. Inspired by turning mechanisms in cockroaches, we demonstrate the potential to achieve turning during dynamic running by varying the stiffness of a single leg. For this work, stiffness was adjusted by physically exchanging the nominal legs for those fabricated from a stiffer material. We also experimentally establish the feasibility of using shape memory alloy wire to actively induce a change in the dynamic stiffness of the leg, providing a path to integration. Finally, preliminary data suggest that a large change in stiffness may enable dynamic maneuvers such as full-speed 90◦ turns over the course of 5 strides.

5.1 Robot Design

A major challenge in building a small-scale light-weight many-legged robot is the minimiza- tion of the number of actuators required to produce effective locomotion. The primary gait used by cockroaches running at high speed is the alternating tripod gait in which ipsilateral front and hind legs move in phase with the contralateral middle leg. Inspired by the incredible locomotory performance of these animals, we sought to design a robot’s kinematics such that all six legs could be driven by a single actuator to produce the alternating tripod gait at high speeds. Though it may seem arbitrary to restrict ourselves to a single motor for forward running, a minimal number of actuators is generally desirable for small-scale robots due to constraints on power density. In fact, for legged robot designs employing one motor per leg, each motor only does positive work on the robot for approximately 30% of the stride.

68 Figure 5.1: The DynaRoACH robot.

We have chosen a brushed DC motor as our primary power actuator. DC motors have a variety of advantages; they are: relatively power dense [83]; inexpensive; straightforward to control; and commercially available in a variety of windings, yielding a spectrum of torque and speed characteristics. However, due to the small size of the motor and therefore high operating speed, we employ a large gear reduction to obtain stride frequencies in a useful range, approximately 8-20Hz.

5.1.1 Kinematic Transmission Design

The primary difﬁculty of using a single DC motor to drive a many-legged robot is the kinematic design of a transmission that couples the angular output of the motor to the oscillatory motion of the legs. The WhegsTMfamily [67][60][18] of robots uses a single motor to power hybrid “wheel-legs” through a standard transmission that uses elements like gears and chains. The iSprawl [55] robot uses a single DC motor with an attached slider crank to drive an arrangement of push-pull cables that lengthen and shorten its legs while protraction and retraction are accomplished passively using compliant rotational hips made from a viscoelastic polymer. In the fabrication of the transmission for this robot, we have chosen to use a prototyping process [44] that is a scaled version of the approach used in smart composite microstructures (SCM) mechanisms. This method has many advantages: it is quite inexpensive; its planar nature enables rapid fabrication and assembly; it produces strong, lightweight mechanisms free of friction and backlash; and the compliant mechanisms used in the robotic design can be synthesized and analyzed using standard kinematic approaches. The robot’s kinematics are comprised of two primitive single degree of freedom mechanisms:

69 Total Mass 23.7 g Body Size 100 x 45 x 30 mm Maximum speed 1.4 m/s 14 body lengths/s Maximum stride 20 Hz frequency Mean Metabolic 3.5 J/(Kg m Cost of Transport · Motor Didel MK07-3.3 Full River 90 mAHr Battery lithium polymer Microcontroller dsPIC33F Bluetooth (Roving Communications Networks rn41 @ 230 Kpbs) Table 5.1: Robot physical parameters

(a) Rear view of the ideal robot kinemat- (b) Side view of the ideal robot kinematics: ab- and adduction of legs from op- ics: protraction and retraction of the legs posite tripods occurs out of phase when is controlled by motion of the middle the middle member of the linkage is member in the fore-aft direction. The translated vertically middle leg moves out of phase with the front and hind legs.

Figure 5.2: Ideal robot kinematics demonstrating the kinematic coupling enabling an alternating tripod gait. The motor output is aligned with the sagittal plane and a crank provides the vertical and fore-aft motion depicted in each ﬁgure.

70 the slider-crank linkage and the parallel fourbar. Rear and side views of the ideal kinematics of the robot are shown in Fig. 5.2. The slider crank linkages enable ab- and adduction of the legs, while fourbar linkages enable protraction and retraction.

Fourbar Kinematics

The fourbar linkages use a simple parallel geometry and therefore have a transmission ratio of unity. In fact, the driving conﬁguration of the fourbars is an inversion of the standard conﬁguration in which the crank is driven and output is taken from the rocker or coupler link. In the case shown in Fig. 5.2b, the coupler is driven by the output of the motor and the output of the fourbar is taken at one of the crank links to which a leg is attached. The fourbar was designed for a nominal protraction angle of 42◦. ± By inspection of Fig. 5.3, the output protraction angle of the leg with respect to the vertical is given by:

1 h θ = sin− (5.1) r f In Eqn. 5.1, h is due to the motion of the motor crank and is simply given by:

h = c sin α (5.2) where c is the length of the crank attached to the motor output and α is the crank angle measured with respect to top dead center.

rf θ

h Figure 5.3: Schematic of fourbar kinematics

Slider Crank Kinematics

In Fig. 5.2a, the left slider crank is a kinematic inversion of the linkage on the right, and the two legs shown belong to opposite tripods. When the right leg is adducted, the left is abducted and vice versa. Each slider crank was designed to give a nominal abduction angle of 40◦ measured from the vertical.

71 To design the slider cranks, we solve the inverse kinematics problem — in our conﬁguration, the slider is the driven element while the output is taken from the crank which functions as a hip (see Fig. 5.4). The constraint equations are deﬁned as follows:

r cos ψ A = s (5.3) r sin ψ s e B = (5.4) s

(B A)T(B A) = L2 (5.5) − − Note that the link r is shared by both the fourbar linkage and the slider crank. However, in the slider crank formulation, rs is the projection of rf onto the transverse plane:

rs = rf cos θ (5.6)

The input from the motor crank drives the slider in the vertical direction giving the relation:

s = s c(1 + sin α) (5.7) 0 −

In Eqn. 5.7 s0 is the initial value for s. Eqns. 5.3- 5.7 are solved at top and bottom dead center of the crank to determine a value for the length of the coupler L that gives the 40◦ abduction angle.

L B

A ϕ

s rs ψ

Figure 5.4: Schematic of slider crank kinematics

Eqns. 5.1- 5.7 can be combined with the following loop closure equations for the slider cranks [62] to solve the 3D forward kinematics:

r cos ψ + L cos φ e = 0 (5.8) s −

72 r sin ψ + L sin φ s = 0 (5.9) s ± The sign of the s term in Eqn 5.9 is determined by whether the slider crank is a member of the left or right tripod. Three projections of the solutions to Eqns. 5.1, 5.8, and 5.9 are shown in Fig. 5.5.

5.1.2 Leg Design

We have chosen a semi-circle C shape for the robot’s leg similar to the design used on the RHex robot [75]. The leg is manufactured by molding using a stiff polyurethane elastomer with 20% softener by weight (PMC-790 rubber and SO-FLEX softener, Smooth-On, Inc.). The C shape offers three primary advantages: lower vertical stiffness, directional collapsibility for obstacle climbing, and a rolling instead of point ground contact. Lower vertical stiffness is obtained by changing the primary loading condition from material compression (as in a straight vertical leg) to material ﬂexure. This allows for another actuator to more easily adjust the overall leg stiffness or to change the length of material allowed to deﬂect in a manner similar to the design in [34]. Collapsibility for obstacle climbing is achieved during forward running, where the open section of the C faces backwards. Upon collision with an object that is taller than the swing clearance height, the C can fold backwards, allowing the leg to continue its forward motion and position some portion of itself on top of the obstacle. During the stride phase, that leg can help pull the robot up and over the object. The round shape of the C along with its compressibility allows it to form a rolling contact with the ground instead of the isolated point contact associated with a typical straight leg. The rolling behavior resembles a distributed foot, creating less vertical displacement of the center of mass and allowing for smoother overall motion [65].

5.1.3 Power, Communication, and Control Hardware

The robot uses the dsPIC33FJ128MC706-based microcontroller board described in [14] for control. Wireless communication is achieved using the Roving Networks rn41 Bluetooth radio interfaced to the PIC microcontroller. Full H-bridge motor control is possible with a custom 2 channel H-bridge motor driver board. However, for this work, the robot was run in a purely open loop, feedforward mode. The motor driver board includes a comparator circuit that enables real- time measurement of the motor back EMF voltage. The analog to digital converter on the PIC is conﬁgured to sample during the off portion of the pulse width modulation signal used to modulate the voltage of the motor drive signal.

73 (a) Top view

(b) Side view

Figure 5.5: Schematic views of the 3D foot trajectories of the robot plotted in the body-ﬁxed coordinate frame.

74 5.2 Models for DynamicSTEERING Turning BY in TRANSIENT Hexapods DESTABILIZATION 275 precisely straight we seta ¯ = 1, ¯b = +1 (dashed CoM path), and to reverse angular momentum − In [52and], Jindrich turn rightand Full we argue seta ¯ that= all2, legs¯b = of +1 a cockroach.5 (dot-dashed can contribute CoM path). to turning These maneuvers. respectively give A =1.618, Inspired byA = measurements 0 and A = of0 leg.615 kinematics for− the turningground reaction strides, forces, after they which devisea ¯ and a simple¯b revert planar to the values 0.5, 1.0 model to describefor which turning.A =0 Then.−25 and they the propose paths a metric rapidly for straighten. leg turning effectiveness In the limited that represents context of the peg-leg− model, the abilitymoving of the leg the to TD generate point forces forward perpendicular along the to body the heading mimics vector extension and apply of the the outside metric front leg observed in to the model.insect While turning the front [24], leg and on it the provides outside of our the first turn evidence was found that to produce foot position the most can force modulate a turn. perpendicular to the heading, the outside middle leg was also found to produce forces that rotate the cockroach appropriately. However,3. their THE model LATERAL requires the LEG middle SPRING leg to produce MODEL additional force parallel to the heading in order to compensate for the resulting change in body orientation relative to theIn heading. this model a single effective leg represents the support tripod of front, rear and contralateral middle legs that many insects, including cockroaches, employ during normal running [3, 4]. A rigid In anbody, effort to again further with explore mass them dynamicsand moment of turning, of inertia ProctorI, and is equipped Holmes demonstrated with two massless that elastic legs that the lateralpivot leg spring freely (LLS) at a model point [76 P], on a low-dimensional the centerline, conservative a distance modeld from of the locomotion CoM (d> in 0 denoting toward the plane,the is capable front of of the producing body). dynamic A left (resp. turning right) [71]. By stride momentarily (stance phase) changing is leg initiated stiffness at TD by placing the and touchdownrelaxed angle, leg, they with demonstrate length l0, at that an translational angle +β kinetic(resp. energyβ) relative can betemporarily to the body ex- axis and allowing the changed forfoot rotational F to pivot kinetic freely energy, about thus the inducing resulting a turn. position, In particular,− which theyremains predict fixed significant until the leg spring is again turning ratesuncompressed, for increasesat in spring which constant point LO of occurs50% and shifts the right in the (resp. location left) of stance the center phase of immediately ensues ≈ pressure (CoP)(a 50% of dutyless than cycle).10% The ofcenter total body of pressure length; results(CoP) fromd is the fixed simulation throughout investigation each stride. See Fig. 8(a), ≈ are reprintedand in refer Fig. to5.7. [7] for further details.

Figure 5.6: A schematic of the lateral leg spring (LLS) model and relevant physical parameters from [71]. Fig. 8. The LLS model showing the body and leg geometry (a) and the coordinate system for the left stance phase, pivoting at F (b). That for right stance is obtained analogously, with TD angle β, as described in the While Proctortext. and Holmes’ model seems to support Jindrich and Full’s assertion that turning− can be achieved with just small deviations from parameters for straight running, one drawback is that LLS parameters cannot be directly3.1. mapped Equations to the actual of Motion parameters andStride of a six-legged to Stride system. Map For example, LLS collapses a tripod of three legs into a single “effective” leg, making it impossible to To describe dynamics during a stride we adopt a polar coordinate systems based at the stance foot. Since the leg is elastic, the system now has three degrees of freedom (two translation and one rotation), described by the generalized coordinates (ζ, ψ) that locate the CoM r(t), and the body angle θ. Fig. 8(b) shows75 the coordinate convention for left stance, with ψ increasing counter-clockwise; for right stance ψ increases clockwise, but θ is taken increasing counter-clockwise throughout. To accommodate this we preface θ with the factor ( 1)n where left and right strides correspond to n even and odd respectively [7]. In terms of these coordinates− the kinetic and potential energies are given in Eqs. (14–15): 1 ˙2 2 ˙2 1 ˙2 T = 2 m(ζ + ζ ψ )+ 2 Iθ , (14) V = V (η), η = ζ2 + d2 +2ζd sin (ψ ( 1)nθ), (15) − − def L = mζ2ψ˙ !+( 1)nIθ˙ = p p = const. (16) F − ψ ± θ REGULAR AND CHAOTIC DYNAMICS Vol. 13 No. 4 2008 278 PROCTOR, HOLMES

Fig. 10. Contour map showing the net angle turned in radians as a function of the turning parameters kt and dt for 3-step turns of the LLS model initiated at left TD, as described in the text. The normal parameter Figure 5.7: Fromvalues [71d =]:0. A002,k contour=0.53 for mapstraight of running turning are at the angle bottom leftin corner radians of the overfigure. 3 strides as a function of the model parameters,− leg spring constant, k, and distance between the center of pressure (CoP) and the center ofFigure mass 9 illustrates (CoM) anin example mm. of the CoM path resulting from a 3-step turn and in Fig. 10 we show how the magnitude of the turn depends on the parameters kt and dt. These results were computed by numerically integrating the equations of motion (17-19), allowing the solution to 1 stabilize on straight running at 0.3 m s− , and then initiating a 3-step turn as described above. Fig. 10 shows that, as expected, the CoP position dt, which controls the moment arm about the differentiate betweenCoM, exerts an the outside greatest influence, front leg and and that an stiff outsideening the leg middle beyond leg the standard because value thosek =0 two.53 legs belong 1 to different tripodsNm− further (but are increases located turning on angle the magnitudes. same side While of rightthe body). turns are best produced by shifting the CoP forward of the CoM at left TD, as described above, substantial left turns result from leaving the CoP at its normal position in left stance and simply reducing d to zero in right stance. For this reason,This deprives we the attempt body of its to usual induce clockwise turning turning of moment a real in right physical stance, leaving robotic the leftwardsystem by directly angular velocity at TD unchanged and thus producing a net turn to the left in each full step. The adjusting legnet stiffness angle turned and per characterizing stride is approximately the turning zero when behaviord = 0.00035, that although results. there By is producing a weak stiffness- t − modulated turningdependence in a on physical,kt. hexapedal robotic model, we aim to increase understanding of the It is notable that turns through 1.5 radians ( 85o) can be obtained by transient shifts in CoP role of leg stiffness and propose a bio-inspired control≈ approach suitable for small, underactuated of less than 3 mm and stiffness increases of 50% (d = 0.002 dt =0.001 and k =0.53 kt = 0.8). Considering that Blaberus has a typical≈ body length− of 2#→ 3 cm, this finding supports#→ the robots. − hypothesis that turning is a minor modification of straight running. (For larger kt and dt values beyond the upper right of Fig. 10, the LLS model rotates through 180o during a single stance phase, exhibiting unrealistic behavior due to the free pivot at the CoP.) Figure 11 shows the changes in kinetic and potential energy that accompany the right turn of Fig. 9. During normal stance phases in straight running before the turn, potential energy rises and 5.3 Experimentalfalls as the leg compresses Setup and uncompresses and both components of kinetic energy fall and rise in phase. Similar qualitative patterns in translational kinetic and potential energy persist through the turn, but rotational kinetic energy increases in each left stance phase of the turning strides, after which it drops to the normal range within one stance phase as the straight gait restabilizes. ExperimentsAs evidenced were byrecorded the increased with peaks high-speed in potential energy video during from the turn, an the overhead outside (left, camera here) leg (250fps, 8bit grayscale, 1280force is1024px, higher than AOS normal Technologies). to provide the netThe turning frame moment sequences and change in were linear downloaded momentum. to a desk- However,× the effect on translational kinetic energy is small, since at normal running speeds it is top computertwo in uncompressed orders of magnitude RAW greater format than typical and rotational processed energies. with the Python programming language Figure 12 shows time histories of system states for the LLS model and for a cockroach during a using OpenCVright [2 turn.] and The the left SciPypanel shows package translational [3]. velocity, The image acceleration, data body from angle the and high-speed heading angle camera was used for 3D reconstruction(direction of the CoM of the velocity robot’s vector), position along with and body orientation angular velocity in the and environment. acceleration over a 3-step turn for the LLS. The right panel shows comparable data (also with the estimated CoM path curvature) for a cockroach, taken from [24]. The conservative LLS model cannot reproduce the insect’s significant reduction in speed as it slows to avoid the obstacle that induces the turn, nor does

5.3.1 Substrate REGULAR AND CHAOTIC DYNAMICS Vol. 13 No. 4 2008

A 1m 1m sheet of 1/8 inch thick cork was ﬁxed to the laboratory ﬂoor, and experiments were × performed entirely on the cork’s surface; running on cork reduces foot slippage. However, the cork also introduces an additional compliance into the system. Based on the experimentally determined value of the Young’s modulus of cork reported in [37], 20 MPa, we can estimate an approximate stiffness for the cork substrate. Assuming a dynamic load of twice the full weight of the robot supported by a single leg contacting over an area of one 1 mm2, the approximate stiffness of the cork is 6500 N/m — more than two orders of magnitude larger than the typical leg compliance.

76 5.3.2 Leg Stiffness

For the turning experiments we adjust leg stiffness by replacing the middle leg fabricated from soft polyurethane rubber containing 20% plasticizer (E 10.7MPa) by weight with a leg of the ≈ same geometry fabricated from the same polyurethane without plasticizer (E 18.2 MPa). Flex- ≈ ural moduli for the two materials were determined experimentally using a standard end-loaded linear beam test.

5.3.3 Camera Calibration

Our goal was to use high speed video data to estimate the kinematic state of the robot during trials in order to better understand its dynamic interaction with the environment. In order to do this estimation from video data, the camera had be calibrated for the volume in which the robot runs. The more accurate the calibration, the more accurately kinematic parameters can be estimated from video data. The calibration process consisted of estimating parameters intrinsic to the camera (such as lens distortion and tangential skew) and extrinsic parameters, or those describing the transformation between world coordinates and camera coordinates.

Intrinsics

Before conducting experiments with the robot, multiple image sequences were recorded of a planar calibration chessboard (7 8 squares, 29mm per side) held in different poses covering the × image plane. We model the camera as a pinhole with fourth order radial lens distortion and zero tangential skew. The intrinsic parameters were estimated by running the software described in [59] on our calibration sequences.

2D Extrinsics

To support estimation of the kinematic state of the running robot, we estimated the equation of the plane containing the cork surface in the camera’s reference frame. To simplify the estimation, we restricted the chessboard to slide across the cork surface for one sequence of calibration images, eliminating out-of-plane rotation and translation. Adding this constraint allowed us to simultaneously estimate the cork plane together with the three-dimensional pose of the chessboard on the surface in each video frame using nonlinear least-squares. This approach was superior to ﬁtting a plane to the 3D point cloud obtained by estimating the chessboard’s pose independently in each sample.

3D Extrinsics

To extract three-dimensional trajectory data, we use a single camera viewing a scene from two perspectives; one view is direct, and the other passes through a slanted mirror. We jointly estimate

77 the intrinsic camera parameters and extrinsic rigid transformation between the stereo pair using the software described in [59].

5.3.4 Tracking Geometry

Retro-reflective spheres (Vicon, Los Angeles, CA) were affixed to laminated strips of balsa wood glued to the robot’s chassis; see Fig. 5.8. Aligning a pair of HMI lights (Super-Sun Gun 200W, Frezzolini) with the camera axis enabled segmentation of these spheres in videos via simple thresholding. The three-dimensional geometry of the spheres affixed to the robot was estimated using the two-camera setup described in Section 5.3.3 using a video of the robot translating through the field of view. The 3D location of each reflective sphere was estimated independently in each sample, then these observations were rigidly transformed to align about the origin. We then aver- aged these 3D locations to obtain an estimate of the geometry.

Figure 5.8: The robot chassis with rigid tracking geometry and retro-reﬂective spheres attached.

5.3.5 State Estimation

Using the calibrated cork surface and robot tracking geometry from Sections 5.3.3 and 5.3.4, it is possible to estimate the robot’s rotations and planar position using pixel observations of the reﬂective spheres from a single overhead camera. We perform state estimation using an Unscented Kalman Filter (UKF) whose state is comprised of Euler rotations (θx, θy, θz) about the x, y, and z axes and planar position (x, y) (for theoretical details about the UKF, refer to [53]; for implementation details, see [81]).

78 5.4 Results

5.4.1 Running Speed

In both straight ahead running and turning, the top speed of the robot was estimated to be 1.4 m/s or 14 bodylengths/s. Fig. 5.9 shows the robot’s running speed as a function of stride frequency for all trials including turning and straight running.

Figure 5.9: Forward speed vs. stride frequency

5.4.2 Cost of Transport

As we noted in the previous chapter, one performance metric for locomotion efﬁciency that’s used among biologists is metabolic cost of transport. Recall that cost of transport is a measure of the total amount of energy required to move one kilogram of body mass one meter, and the millirobot, RoACH, has a “metabolic” cost of transport of more than 18,000 J/Kg m. − To measure the metabolic cost of transport on DynaRoACH we measure the battery voltage and the back EMF voltage while the robot is running. The difference of the two combined with the motor resistance determines the electrical power supplied to the motor simply from Ohm’s law:

2 (Vbatt Vemf ) Pelectrical = − (5.10) Rmotor The electrical power consumed is divided by the product of the mass of the robot and its speed (as estimated from high speed video using the UKF) with the result being “metabolic” cost of transport. For the DynaRoACH robot, a plot of cost of transport at a variety of running speeds is shown in Fig. 5.10. Though it’s biased by a larger number of samples at higher stride frequencies, we note that the mean metabolic cost of transport for the robot is around 3.5. This value is signiﬁ- cantly lower than many other values reported in the literature recently [15] for small legged robots. Some possible explanations for such a low value include the very low-loss compliant transmission and well-tuned leg stiffness enabling dynamic energy storage and recovery.

79 Figure 5.10: Metabolic cost of transport vs. stride frequency for the DynaRoACH robot.

5.4.3 Turning by Stiffness Control

Results from the study of turning in cockroaches suggest that all legs have the potential to contribute to dynamic turning. Small kinematic changes to the front leg were observed during turning maneuvers, while additional force production was required from the middle and hind legs to explain turning according to the simple dynamic model proposed in [52]. With the biological results in mind, we chose to adjust the stiffness of the middle leg in each tripod in order to affect turning. By increasing the stiffness of the middle leg, we reason that we can generate larger forces in that leg and induce a yaw torque that will cause the robot to turn. With the stiffer leg in place, the robot was run over a smooth surface at a variety of speeds. The results from the turning experiments are shown in Figs. 5.11 and 5.12. Fig. 5.11 demonstrates that by stiffening the outside leg, we are able to consistently achieve turning in either direction. Qualitative examination of the ﬁgure, however, reveals that the robot turns more smoothly to the left than to the right, a fact that may be explained by small kinematic asymmetries in the robot’s gait. Likewise, Fig. 5.12 shows that the maximum turning rate is relatively insensitive to stride frequency while there appears to be a wide scatter of turning rates possibly due to as yet unmodeled frequency-dependent dynamics. Data for the left turn show a mean turning rate of approximately 75 degrees per second, while the mean turning rate for right turns is approximately 50 degrees per second with signiﬁcantly more variation in the data.

80 Figure 5.11: Summary data from all turning and straight running trials at 8-20Hz stride frequencies.

Figure 5.12: Turning rate vs. stride frequency

81 SMA wire actuator

Figure 5.13: A CAD depiction of the compliant semi-circular leg with shape memory alloy wires running along the outside of the leg.

Thursday, April 8, 2010

5.5 Adjustable Stiffness Leg Design

To actively initiate turning, we propose adjusting the stiffness of the middle leg on each side using shape memory alloy wire attached to the leg as shown in the CAD drawing in Fig. 5.13. The wire is 75 µm diameter Flexinol (Dynalloy, Tustin, California, USA) that is run from the hip down along the outside of the C, then turned around on a piece of plastic embedded in the leg material before running up along the C back to the hip, where it is connected to control electronics. Actu- ation of the SMA wire increases its modulus and decreases its length. The result is a significant increase in the area moment of inertia and, thus, the flexural rigidity of the whole leg. To verify that our design can achieve the same increase in stiffness as fabricating the legs without plasticizer, we tested the dynamic stiffness of a “soft” leg with an actuator attached using a muscle lever (300C, Aurora Scientific) to apply varying loads at frequencies in the range of the running frequencies the robot is capable of. Fig. 5.14 shows that at applied dynamic loads between 50mN and 80mN, the actuated leg is capable of matching or nearly matching the stiffness of the harder polyurethane. We are encouraged by this result as it demonstrates a clear integration path for active leg stiffness modulation.

5.6 Discussion

We demonstrated turning and rapid running in a small underactuated legged robot. Simply by stiffening the middle leg on the opposite side of the desired turn direction, we induced turning at rates of roughly 50 degrees per second to the right and 75 degrees per second to the left. At certain “troubling” stride frequencies, the robot dynamics were less predictable and hence turning

80 50mN stiff, passive 80mN stiff, passive 60 100mN stiff, passive 50mN soft, unactuated 40 80mN soft, unactuated 50mN soft, actuated 80mN soft, actuated 20 100mN soft, actuated Sti ff ness (N/m)

0 1 5 9 13 17 21 Frequency (Hz)

Figure 5.14: Experimental measurements of dynamic leg stiffness comparing a passive, “stiff” leg to a “soft” leg with an integrated shape memory alloy (SMA) wire actuator with applied dynamic loads of 50mN, 80mN, and 100mN.

behavior less repeatable; we believe that further kinematic refinements will help to reduce the frequency dependence. Based on simple planar biological models for turning in hexapods, we anticipated that the middle leg would initiate turning. Although the model in [52] predicted that increasing the front leg stiffness would be most effective in causing turning, the middle leg gave better results for our robot. This inconsistency may arise simply due to differences in gait kinematics and scale between our robot and the cockroaches studied in [52]. In particular, our robot’s limbs are significantly less sprawled than a cockroach’s, possibly explaining the relatively large out-of-plane rotational dynamics reported in Tables 5.2 and 5.3. These dynamics, however, are not modeled in [52] or [71]. Fig. 5.15 shows the evolution of the pitch and roll rotations as a function of phase for a characteristic trial in which the robot turned as well as one in which it ran straight. For the turning trial, roll and pitch clearly oscillate antiphase, whereas, for straight-ahead trials, they appear to generally oscillate in phase. Consistent with aggregate data in Tables 5.2 and 5.3, the roll and pitch amplitudes are smaller for straight running. Altogether, these observations suggest that roll and pitch likely play a significant roll in the control of turning by adjusting leg stiffness. At specific stride frequencies, we observed rapid dynamic turns in which the robot turned ≈ 90◦ over roughly 5 strides; see Fig. 5.17 for data from an illustrative trial. Of note is the fact that large instabilities in the robot’s roll (rotation about the major body axis in the horizontal plane) and pitch (rotation about the body minor axis) correlate with the turns, settling out immediately following the execution of the maneuver. We believe this is an exciting result deserving of more

83 (a) (b)

Figure 5.15: Roll and pitch angle vs. phase for characteristic trials of (a) straight running at 13.5 Hz and (b) turning at 15 Hz

Figure 5.16: Angular orientation for the robot during dynamic turns. Note increases in roll and pitch magnitudes corresponding to sudden changes in yaw.

84 Stride Frequency Min. Roll Max. Roll Min. Pitch Max. Pitch 10-15Hz -7 3◦ 9 4◦ -6 4◦ 7 3◦ ± ± ± ± Table 5.2: Roll and pitch angle statistics for n=14 straight-running trials

Stride Frequency Min. Roll Max. Roll Min. Pitch Max. Pitch 5-10 Hz -14 3◦ 13 2◦ -10 2◦ 11 2◦ ± ± ± ± 10-15Hz -10 6◦ 10 7◦ -9 4◦ 8 4◦ ± ± ± ± 15-20Hz -10 8◦ 9 6◦ -10 5◦ 10 5◦ ± ± ± ± Table 5.3: Roll and pitch angle statistics for n=31 turning trials

study; the ability of a legged robot to reliably execute such sharp turns could enable extremely rapid navigation.

Figure 5.17: Position data for the robot during a series of dynamic turns. The time interval between markers is 2.5ms and the stride frequency is approximately 13 Hz.

Our demonstration of consistent turning via modulation of leg stiffness has implications for biomechanics research. Since it is extremely easy to adjust leg stiffness by simply replacing the leg, the robot can support systematic studies of the parametric dependence of maneuverability. We have already provided preliminary experimental validation for the turning predictions from simple models like LLS, since we obtain substantial turning rates using the predicted scales of stiffness adjustment. One interesting future direction is to extend this result by carefully characterizing the connection between the lumped-element parameters in (for example) LLS and the design parameters of the physical robot.

85 Part IV

Conclusions and Appendices

86 Chapter 6

Conclusions

This work has explored the potential for designing legged robots at the milli- and meso-scales using the folded compliant mechanisms produced by the smart composite microstructures process and its analog scaled prototyping process. We’ve developed a spatial, strain energy-based approach to modeling and assessing the stiffness properties of parallel compliant mechanisms and experimentally demonstrated its ability to improve our designs through kinematic redundancy. We have also designed and built two autonomous legged robots at the milli- and meso-scales using SCM and scaled prototyping, respectively. To our knowledge, the 2.4g RoACH robot is the smallest legged robot to date capable of untethered, sustained, steerable locomotion. With the 24g DynaRoACH robot we have reduced actuation requirements to a single DC motor and improved efﬁciency through dynamic operation. The minimally actuated DynaRoACH design has also demonstrated the potential for controlled turning as well as highly dynamic maneuvers using the biologically inspired approach of modulating leg stiffness.

6.1 Future Work

However, even with the accomplishments in this dissertation, we have only begun to scratch the surface of what is possible with SCM (and its scaled prototyping analog) and power-dense, small-scale legged robots. Much work remains to be done to improve the capabilities of these robots as well as the tools used to design and fabricate them.

6.1.1 Improved Design Tools

Folded SCM mechanisms are a relatively new approach to designing and fabricating robots at the milli- and meso-scales. Their combination of kinematic mechanism design, compliant mechanism theory, composite materials, and planar fabrication make them difficult to model in a unified way using existing computer-aided design (CAD) tools. Linkage synthesis and analysis software can be helpful in the kinematic design process, and finite element software is useful for ensuring that structural loads don’t exceed maximum allowable stresses. Where fabrication is concerned,

87 two-dimensional representations suitable for manufacturing on a laser-machining system can be produced using any number of drafting packages. However, no ability exists yet for the folded robot designer to synthesize a three-dimensional design directly, analyze it structurally or kinematically, and generate an unfolding that can then be manufactured automatically. It is this manual translation of a folded design to a suitable 2D representation that is error prone and necessitates the scaled prototyping process in AppendixA. In the absence of effective computer aided design tools for these robots, the best tool for design evaluation is still a functional prototype that can be realized in less than an hour.

6.1.2 Integration of Electronics into Robot Structure

Printed circuit board and folded robot fabrication share a very important property - they are both planar. Integration of electronics with our existing folded robots is almost an afterthought in its current state. As robots get smaller and more dynamic, their performance becomes much more sensitive to mass distribution. Having the ability to integrate electronics (as the Microrobotics group at Harvard has begun to do [23]) and potentially even batteries into the robot structure would greatly simplify design and improve overall robustness. With integrated electronics, robots could be more easily ruggedized for operation in harsh or aquatic environments if necessary.

6.1.3 Dynamic Characterization and System ID

Clearly, from the results in Chapter5, a great deal of interesting behavior is embedded in the dynamics of the DynaRoACH robot. Though we have proven the viability and capability of the design and shown the potential for turning control with small changes to leg stiffness, the full underactuated dynamics are not yet well understood. Reduced order models like SLIP [16] and LLS [76] are useful for providing insight into the dynamic behavior of animals and potentially robots, but mapping predictions from the model to the vastly more complicated physical system is difficult if not currently impossible. One interesting approach to improving our understanding of these dynamic systems is the no- tion of “fitting” their dynamics to existing models in order to close the loop between model and design parameters like leg stiffness, stride frequency, leg morphology, etc. This is a very challenging problem, and its solution is far from clear. A combination of experimental techniques from biology like 3D motion capture and simultaneous ground reaction force measurements combined with engineering techniques from system identification could begin to shed light on the underlying dynamics of these systems.

6.1.4 Dynamic Climbing

The high power density of small SCM-based robots makes them ideal for challenging locomotion modes like dynamic vertical or inverted climbing. With a better understanding of their dynamics and current advances in adhesion technology [56][61], legged millirobots could make

88 very effective, dynamic climbing platforms opening up a large space of applications in exploration, search and rescue, and structural monitoring.

6.1.5 Robots as Platforms for Testing Biological Hypotheses

Finally, perhaps one of the most exciting applications of these small SCM robots is at the interface of integrative biology and robotics. The ease with which robot designs can be modiﬁed and fabricated makes the scaled SCM process ideal for quickly generating a variety of morphologies to help test hypotheses about animal locomotion performance. While this dissertation has hopefully made some progress toward the goal of realizing capable legged locomotion at the scale of a cockroach, we are clearly still very far from being able to produce the truly awesome locomotion capabilities that we see in these remarkable insects. We can only hope that some ideas contained in this work will inspire current and future engineers and biologists to take up the mantle and produce new approaches that will help us create insect-like locomotion in robotic platforms.

89 Appendix A

A Scaled Prototyping Process for Folded Millirobots

In this appendix we introduce an exact scaled version of the smart composite microstructure (SCM) process [91] that enables the millirobot designer to produce functional scaled prototypes of a robot design. The process utilizes very low cost, so-called “green” materials and allows for the physical realization of a robot prototype in less than one hour for the basic structure without wiring or actuators at a cost of less than $1 per robot. While the material costs for an SCM robot are similar, the speed of the SCM process is fundamentally limited by composite cure times, laser micro-machining times, and difficult hand assembly of very small structures. However, by closely mimicking the SCM process, we enable the designer to produce accurate macro models at anywhere from a 2-10 scale so that kinematic and static loading conditions can be explored × quickly with a hands-on model. The speed and negligible material cost of the process significantly shorten the design cycle for an SCM-based millirobot. For an overview of the SCM process we refer the reader to Chapter3. In the following sections, we enumerate the steps necessary to produce scaled prototypes, briefly review design rules adapted from the SCM process for folded, flexure-based robots, and discuss the effects of scaling on the overall performance of a prototype.

A.1 Process Overview

A.1.1 The Prototyping Process

The macro-scale prototyping process closely follows the SCM process, but instead of pre- impregnated composite fiber laminate, we substitute lightweight, double sided coated posterboard (Nature Saver posterboard, available from Officemax.com). For the polyimide flexure layer, we substitute a thicker polyethylene terephalate (PET) film because thermal compatibility of the film is less of a concern than in the SCM process. In addition, the PET is less inert so it bonds better to various adhesive polymers. Just as in the SCM process, the posterboard is cut using a precision

90 laser (VersaLaser VL200) according to a 2-dimensional design drawing. However, unlike with carbon ﬁber composites, the posterboard can be easily cut with an infrared laser. The steps for the prototyping process are outlined below with the step numbers corresponding to Fig. A.2:

Figure A.1: Steps for prototyping a hexapod crawling robot: (1) A 2-dimensional drawing is made in a vector-based illustration program (Solidworks). (2) A blank piece of posterboard with a pre- cut fold line is folded and placed in a laser cutter, and the ﬂexure cuts are made in both layers. (3) The workpiece is removed and adhesive is applied to both sides inside the fold. A 50 µm ﬁlm of PET is sandwiched between the two sides of the fold. (4) The workpiece is fed through heated rollers at 230◦F at a low speed (approx. 5 mm/sec). (5) The workpiece is placed back in the laser cutter and outlines of the parts are cut out. (6) Parts are released from the workpiece. (7) Individual linkages can be folded up in place. The three joints on the two leftmost pieces are fourbars used as hips. The joints in the rightmost piece form two Sarrus linkages when two are folded and glued out of plane. (8) The three plates are joined such that the hips for one tripod sit above the central plate while the hips for the other sit below, and the legs are attached to the hip fourbars. (9) SMA actuators are attached and wired in place and spherical PDMS feet are added.

1. The process begins with a sheet of posterboard. 2. Gaps are laser-cut into the posterboard where flexures will be in the final structure. 3. Two layers of adhesive polymer with one layer of flexural polymer (high melting point) are placed between two sheets of posterboard.

91 (1)

(2)

(3)

Heated Rollers (4)

(5)

Figure A.2: A step-by-step illustration of the posterboard link and hinge fabrication process

4. The resulting structure is rolled through heated rollers to apply pressure and melt the adhesive layers, bonding the structure.

5. Outlines of parts are then laser cut, releasing the parts from the original sheet.

Because the posterboard itself, unlike pre-impregnated composites, contains no adhesive, a simple thermoset polymer film for the flexure layer is insufficient. We must either add adhesive layers separately, or use a polymer film with an integrated adhesive to bond the structure. A carrier film with integrated adhesive is convenient because it’s essentially monolithic. However, the standard thicknesses of such films limit the ability of the designer to change the stiffness of flexure hinges by using different thickness films. While it is possible to use more than one layer of such a film, the interaction of the adhesive layers between the carrier film layers creates a noticeable viscoelastic response from the flexure hinges. In order to retain the flexibility for the designer in the choice of flexure thickness, we have chosen to add a separate adhesive between the posterboard and the flexure film. For the the flexure film we use 50 or 75µm PET depending on the desired flexure stiffness. For the adhesive layer we use either UHU TMglue stick which is a gel- based polyacrylate adhesive, or we pre-laminate the inner faces of the workpieces with hot-mount thermal adhesive. This prototyping method is appropriate for fabricating entire functional robots or simply mock- ing up a small mechanism to be integrated into a larger robot design. The hexapod crawler design presented in Chapter4 highlights the difficulty of visualizing a complicated 3D mechanism that

92 results from folding a 2D design. The posterboard prototyping process enables the designer to quickly try out a folded design and explore features such as alignment tabs or self-jamming flexures that can aid significantly in the assembly process. In addition, because the goal of the prototyping process is to produce a large-scale model of a robot that will ultimately be fabricated using anisotropic, directional composites, the prototype can be used to explore folding approaches that ensure that the composite fibers are properly oriented for the loading conditions that will be imposed on the robot. For example, fiber direction should be perpendicular to any movable flexures because the composite laminate is very weak in bending about the fiber axis compared to bending about the axis transverse to fiber direction.

A.1.2 Flexure Design

It is important to address the issue of scaling in order to effectively use the prototypes for predicting mechanism behavior at the small scale. Ideally we would like to develop a scaling law that can be used to prototype at any scale. There are two parameters in which we’re interested - the flexure hinge behavior and the stiffness of the structural members or links. Simple design rules based on modeling flexures using the pseudo-rigid-body model presented in [47] are outlined in [91]. In short, to mostly closely match the behavior of an ideal pin joint, the ratio of the axial stiffness of the flexure to its rotational stiffness should be maximized. This objective is subject to geometric constraints imposed by the maximum elastic strain sustainable in the flexures as well as the maximum angle of rotation of the flexure before the edges of the rigid links on opposite sides make contact and jam given by:

l φ = (A.1) max t Eqn. A.1 is an important constraint when using the fast scale prototyping process. One limitation of the process is that the posterboard is available only in discrete thicknesses. The material used in our prototypes is approximately 6 times thicker than the true scale material. The implication is that this constraint will be stronger at the larger scale. Thus, we must take care in using the design equations from [91] to design the large scale ﬂexures.

A.1.3 Compensation for Laser Dimensions

In some cases it is important to have an accurate prediction of joint stiffness and range of motion. For example, when specifying required force to actuate a linkage built using the scaled prototyping process, accurate knowledge of individual joint stiffness allows the designer to cal- culate an effective mechanism stiffness. Likewise, some designs require so-called “sacrificial” flexures that bend exactly to a specific angle and are then fixed with a permanent adhesive. This approach is very convenient for creating desired geometry. In both cases, it’s important for the designer to know the relationship between drawn dimensions and the true resultant size of the part. Because the laser beam used to cut the cardboard has some fixed dimension, design drawings must compensate for its size. In addition, because of beam divergence the sidewalls of laser cuts are not

93 vertical as shown in Fig. A.3, and this effect can also have an impact on design functionality. Fig. A.4 shows the relationship between the true flexure length and its length dimension in the design drawing, while Fig. A.5 shows the relationship between true flexure length and angle at which the flexure “jams” or no longer admits rotation.

Adhesive Posterboard Posterboard Adhesive

Flexure Flexure (a) A ﬂexure in which the initial cuts (b) A ﬂexure in which the initial cuts were performed with inner adhesive were performed with the inner adhesive faces down. faces up.

Figure A.3: Effect of workpiece orientation on ﬂexure geometry. Cut sidewall angles are exagger- ated for effect.

Figure A.4: Relationship between the ﬂexure cut length in the design drawing and the resulting artifact. Posterboard was cut with thermal adhesive liner paper face up (PU) and face down (PD).

Linear regressions were performed on both the ﬂexure length and jamming angle data in Figs. A.4 and A.5 to provide the prototype designer with a simple calculation to estimate true ﬂexure dimensions and jamming angles.

94 Figure A.5: Flexure jamming angle as a function of the true ﬂexure length for 440 µm thick posterboard with a 25 µm ﬂexure layer.

A.1.4 Developing a Scaling Law

In order to scale the rigid members of a design, we must formulate a way of relating the material properties and geometry of a structure built at the small scale to those of the scaled structure. Because the materials used at the different scales are so different, we propose a dimensionless parameter to relate the actual size structure to the scaled prototype as follows: kl ξ = c (A.2) g m · where k is the stiffness of a characteristic link in bending, lc is a characteristic length, and mg is the weight of the entire structure. This formulation comes from using structural stiffness vs. weight as a performance metric. This dimensionless parameter is also based on the implicit assumption that the robot is operating under quasi-static conditions. For a dynamic scenario, it would be preferable to use a characteristic load rather than the weight of the robot in the denominator. It is our aim to use this ratio as a guide in designing scaled prototypes of folded millirobots. That is, robotic structures at the macro scale will perform similarly to their true scale counterparts in terms of stiffness-to-weight ratio if they have similar values for ξ. In following section, we compare the values of ξ for both the scaled prototype and the milli-scale 2.4g hexapedal robot design in Chapter 4.

A.2 Comparison of Scaled Prototype to RoACH Millirobot

In this section, we use the example of the hexapod crawling robot from Chapter4 to demonstrate the prototyping process and verify our scaling law.

95 Table A.1: Comparison of a 2 posterboard prototype to S-2 glass ﬁber composite using Eqn. A.2 ×

Posterboard S-2 Glass lc 22 mm 11 mm b (width) 5 mm 2.5 mm h (thickness) 0.875 mm 0.15 mm Papplied 0.01 N 0.01 N 6 6 δobserved 80 x 10− m 665 x 10− m N N Kbending 125 m 66.5 m 2 3 m g 4.6 x 10− N 3.2 x 10− N · Kg Kg ρ 750 m3 2400 m3 E(11) 30 GPa 43 GPa ξ 84.5 61.1

Based on the two robot skeletons depicted in Fig. A.6 we can begin to compare the 2 proto- × type to the true scale version using the scaling law proposed above. The characteristic link we will use to determine the stiffness, k, is the leg link. We will use the leg length as our characteristic length, lc. To determine the bending stiffness of each leg (posterboard and composite) the leg was clamped at one end and subjected to a known load at the free end. The tip displacement was observed with a video camera and measured on screen at a zoom of approximately 32 . Using these × measurements we can compute the value of ξ for both the 2 scale prototype and the composite × ﬁber structure. The measurements and results are summarized in Table A.1. The value of ξ for the 2 prototype is within 30% of the value for the composite ﬁber structure. × In fact, the value of ξ for the 2 prototype is higher than for the true scale model. This indicates that × the prototype may, in fact, be overpredicting the performance of the composite. This discrepancy can be at least partly explained by the fact that the elastic modulus of paper used in the prototypes has been estimated to be on the same order as that for the S-2 glass [69] - approximately 30GPa. In addition, the posterboard is a factor of 3 less dense than the S-2 glass. To even more closely match the prototype to the true scale we could reduce the thickness of the leg links or increase the scaling factor slightly to approximately 2.4 . ×

96 (a) 2 posterboard crawler model ×

(b) True scale S-2 glass crawler structure

Figure A.6: The scaled prototype structure with actuators integrated and the actual size crawler structure (without actuators or wiring)

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