Design and Miniaturization of an Ultra-Fine Multi-Degree-of-Freedom Robotic Instrument for Ophthalmic Minimally Invasive Microsurgery

Design and Miniaturization of an Ultra-Fine Multi-Degree-of-Freedom Robotic Instrument for Ophthalmic Minimally Invasive Microsurgery | Research Square window.SnipcartSettings = { analytics: { enabled: false } }; (function() { var accessVector = localStorage.getItem('access_vector') || ''; window.dataLayer = window.dataLayer || []; if (accessVector) { window.dataLayer.push({ user: { profile: { profileInfo: { snid: accessVector } } } }); } })(); (function(w,d,s,l,i){w[l]=w[l]||[];w[l].push({'gtm.start':new Date().getTime(),event:'gtm.js'});var f=d.getElementsByTagName(s)[0],j=d.createElement(s),dl=l!='dataLayer'?'&l='+l:'';j.async=true;j.src='https://www.googletagmanager.com/gtm.js?id='+i+dl;f.parentNode.insertBefore(j,f);})(window,document,'script','dataLayer','GTM-K279D39R'); Browse Preprints In Review Journals COVID-19 Preprints AJE Video Bytes Research Tools Research Promotion AJE Professional Editing AJE Rubriq About Preprint Platform In Review Editorial Policies Our Team Advisory Board Help Center Sign In Submit a Preprint Cite Share Download PDF Research Article Design and Miniaturization of an Ultra-Fine Multi-Degree-of-Freedom Robotic Instrument for Ophthalmic Minimally Invasive Microsurgery Makoto Jinno, Ryosuke Nonoyama, Iulian Iordachita This is a preprint; it has not been peer reviewed by a journal. https://doi.org/ 10.21203/rs.3.rs-8771621/v1 This work is licensed under a CC BY 4.0 License Status: Under Review Version 1 posted 9 You are reading this latest preprint version Abstract Minimally invasive surgery (MIS) has transformed surgical practice by reducing patient trauma and improving postoperative outcomes. In laparoscopic surgery, these benefits have been further enhanced by the clinical adoption of teleoperated robotic systems, most notably the da Vinci Surgical System, which provides improved dexterity, motion scaling, and ergonomics in confined environments. As surgical robotics advances, its application is expected to extend beyond conventional MIS to microsurgical procedures requiring levels of precision and stability beyond those achievable manually. However, the clinical adoption of robotic assistance in microsurgery remains limited, particularly for minimally invasive procedures in highly constrained workspaces. Teleoperated leader–follower robotic architectures offer a promising solution for robot-assisted minimally invasive microsurgery (MIMS) by enabling precise motion scaling and tremor suppression while preserving intuitive surgeon control. Ophthalmic MIMS requires dexterous manipulation within an extremely confined intraocular workspace under millinewton-level interaction forces. Although snake-like and continuum instruments have been explored to improve access and distal dexterity, achieving multi-degree-of-freedom (DOF) motion within a submillimeter outer diameter remains challenging. These challenges stem from inherent trade-offs among bending range, shaft stiffness, wire routing, pretension, and buckling stability. This work presents the design and miniaturization of an ultra-fine multi-DOF robotic instrument for vitreoretinal surgery. The proposed instrument integrates 2-DOF distal bending (pitch and yaw), shaft rotation (roll), and a microgripper within a 0.7 mm outer diameter. To support miniaturization while maintaining manufacturability and structural integrity, a novel surface-constrained, V-type disk-stacked bending mechanism is introduced. Wire passability through reduced-diameter guide holes is geometrically verified at maximum disk tilt, and shaft stiffness and Euler buckling are analyzed using second-moment-of-area models under a conservative 10 mN lateral tip load. Prototypes with outer diameters of 0.9 mm and 0.7 mm were fabricated and tested. The 0.7 mm instrument demonstrated smooth pitch–yaw bending and reliable grasping, with bending hysteresis of approximately ± 5°. Shaft deflection during bending and grasping remained below 0.06 mm, while rotational whirling produced displacement amplitudes of 0.1–0.23 mm. These results highlight key design trade-offs and provide experimentally validated guidelines for the development of ultra-fine robotic instruments for ophthalmic MIMS. Medical robotics Retinal microsurgery Ultra-fine robotic instrument Disk-stacked bending mechanism Mechanical design Figures Figure 1 Figure 2 Figure 3 Figure 4 Figure 5 Figure 6 Figure 7 Figure 8 Figure 9 Figure 10 Figure 11 Figure 12 Figure 13 Figure 14 Figure 15 Figure 16 Introduction MIS has fundamentally reshaped modern surgical practice by reducing patient trauma and enhancing postoperative quality of life. These benefits have been further amplified in laparoscopic surgery through the widespread clinical adoption of teleoperated robotic systems, most notably the da Vinci Surgical System, which offers superior dexterity, motion scaling, and ergonomic control within confined operative spaces [1,2]. As surgical robotics continues to evolve, there is an increasing expectation that its application will extend beyond conventional MIS toward microsurgical domains, where the demands for precision, stability, and dexterity often exceed the capabilities of manual techniques [3]. Despite this progress, the clinical translation of robotic assistance to microsurgery remains uneven and limited in routine practice. Although several robotic systems have been effectively implemented for microsurgical tasks in open environments [4,5], minimally invasive robotic systems designed to function within highly constrained spaces have not yet achieved widespread clinical adoption. Within this landscape, teleoperated leader–follower robotic architectures offer a compelling solution for robot-assisted MIMS, as they enable precise motion scaling and tremor suppression while preserving intuitive surgeon control. Among microsurgical procedures, vitreoretinal surgery represents one of the most demanding applications in terms of invasiveness and precision. Accordingly, a wide range of robotic systems has been developed to assist retinal surgery [6], with feasibility demonstrated through in vivo animal experiments [7–9] as well as first-in-human clinical studies [10,11]. However, many of these systems rely on straight, rigid instruments, which inherently restrict achievable approach angles and limit dexterity within the confined intraocular space. To address these limitations, snake-like and continuum robotic instruments have attracted increasing attention for retinal surgery. By providing distal bending capability, such instruments enable more favorable approach angles to anterior retinal regions and facilitate technically challenging procedures, including retinal vein cannulation and membrane peeling. Consequently, several bending-type robotic instruments, including snake-like and continuum designs, have been proposed for vitreoretinal applications [12–14], underscoring the potential of articulated and continuously deformable distal mechanisms for enhancing dexterity within the confined intraocular workspace. Joint mechanisms for robotic surgical instruments can be broadly classified into hinge-based articulated mechanisms and continuum bending mechanisms that undergo continuous deformation into an arc shape. For laparoscopic surgical instruments with outer diameters of approximately 7 mm, hinge-based joint mechanisms are commonly employed. However, when accounting for the strength requirements at hinge joints and the associated fabrication complexity, realizing submillimeter-scale instruments with outer diameters below 1 mm using hinge-based designs becomes technically impractical. For MIMS, bending-type mechanisms with simpler structural configurations and component geometries are generally more suitable for miniaturization and diameter reduction. Bending-type mechanisms can be further categorized into three principal approaches. The first is the notched-tube type, wherein a tubular structure is locally notched to create flexible hinge regions and actuated by tendons or wires [13–15]. The second is the concentric-tube type, where multiple pre-curved, flexible concentric tubes are bent through relative rotation and translation [12,16–17]. The third is the stacked-element type, in which discrete elements such as rings, vertebrae, or disks are stacked and bent via tendon or wire actuation [18–21]. Although notched-tube-type mechanisms have been realized with outer diameters on the order of 1 mm, their clinical applicability remains extremely limited, and further diameter reduction is considered challenging due to fabrication constraints and insufficient structural strength. Concentric-tube-type mechanisms have achieved submillimeter-scale outer diameters; however, their large intrinsic curvature and structural characteristics make precise orientation control of end-effectors difficult within highly confined spaces. Consequently, achieving both extreme miniaturization and compact, controllable bending behavior remains challenging for both approaches. In contrast, stacked-element-type mechanisms are considered more amenable to miniaturization while preserving manufacturability, assembly feasibility, and bending performance through appropriate geometric design of the constituent elements. Based on the principle of the variable neutral-line mechanism [21], we therefore investigate a disk-stacked bending mechanism as an instrument unit for robot-assisted MIMS. In our previous studies, we developed an improved robotic intraocular snake and analyzed the kinematics and drive mechanisms of its highly dexterous distal unit, demonstrating the feasibility of an ultra-fine, disk-stacked 2-DOF bending mechanism with an outer diameter of 0.9 mm suitable for intraocular manipulation [22,23] (see Fig. 1(1) and (2)). We subsequently developed a microgripper compatible with this ultra-fine bending mechanism and demonstrated stable grasping capability even within an extremely constrained workspace [24] (see Fig. 1(3)). These studies confirmed that ultra-miniaturized bending mechanisms and distal end-effectors can be successfully implemented for intraocular robotic applications. In addition to the 2-DOF bending mechanism, and with future clinical applications in mind, we also proposed multi-DOF bending mechanisms and corresponding drive systems designed for integration with follower arms of teleoperated surgical robots [25]. However, for practical intraocular surgical applications, outer diameters of 0.9 mm inherently limit the range of feasible procedures, making further miniaturization necessary. Such diameter reduction is expected to introduce severe trade-offs among bending capability, shaft stiffness, wire routing, pretension, and buckling stability. To address these challenges, this study aims to present the design and miniaturization of an ultra-fine multi-DOF robotic instrument specifically intended for ophthalmic MIMS. The proposed instrument integrates a 2-DOF bending mechanism, a shaft rotation axis, and a microgripper within an outer diameter of 0.7 mm. To enhance manufacturability and assembly feasibility at this scale, a novel disk-stacked bending mechanism based on a surface-constrained disk geometry is introduced. In addition, the influences of shaft dimensions, wire routing configuration, and initial wire pretension on bending performance, distal tip deflection, and buckling behavior are systematically investigated through experiments using full-scale prototypes. Through these studies, this paper elucidates the key design trade-offs associated with ultra-fine robotic instruments and provides experimentally validated and practically applicable design guidelines for future ophthalmic MIMS systems. Methods (Mechanical Design) This section discusses the design and miniaturization of 3-DOF robotic instruments with a microgripper. Here, “3-DOF” refers exclusively to distal orientation (pitch, yaw, roll) and does not include insertion/translation, remote center of motion (RCM), or end-effector actuation. Robotic assistance in surgery can be broadly categorized into (i) handheld devices, (ii) teleoperated (leader–follower) robotic systems, and (iii) human–robot cooperative control systems [26]. In laparoscopic surgery, teleoperated systems exemplified by the da Vinci platform have become the de facto standard in clinical practice due to their superior ergonomics, dexterity, and stable manipulation in the abdominal cavity. In minimally invasive microsurgery (MIMS), handheld, teleoperated, and cooperative control approaches are all conceivable; however, as clinical demands expand toward complex and prolonged tasks requiring high reproducibility, teleoperated systems are expected to become increasingly important. Moreover, considering future developments toward automation using AI-based perception and decision-making, the follower robot technology of teleoperated systems provides a natural basis for semiautonomous or supervised-autonomous microsurgical functions. A typical follower robot for teleoperated MIMS consists of (1) a positioning arm providing an RCM constraint to maintain the insertion point (sclerotomy), and (2) a robotic instrument that provides distal orientation DOFs plus an end-effector. In this study, we focus on the design of the robotic instrument that integrates “3-DOF distal orientation + end-effector,” where the distal orientation is realized by a 2-DOF bending mechanism and a shaft rotation axis, and the end-effector is a microgripper. Design Requirements Derived from Ophthalmic MIMS Ophthalmic MIMS, such as vitreoretinal surgery, imposes stringent constraints on instrument diameter, dexterity, and mechanical stability. First, the insertion diameter must be sufficiently small to be compatible with standard trocar-based workflows. While an outer diameter (OD) of 0.9 mm can be used in certain experimental or limited clinical scenarios, its applicability is restricted because it may require enlargement of the sclerotomy or reduce compatibility with standard tool exchanges. Reducing the OD to approximately 0.7 mm, corresponding to 22-gauge-class instrumentation, substantially broadens the range of potential ophthalmic procedures and facilitates integration into routine workflows. Second, the instrument must provide adequate distal dexterity to achieve favorable approach angles and stable manipulation inside the eyeball. For this purpose, we target a distal orientation structure that combines (i) 2-DOF bending (pitch and yaw) and (ii) a shaft rotation axis (roll) while maintaining an end-effector actuation DOF. Third, mechanical stability under wire-driven actuation is essential. In ultra-fine instruments, internal forces generated by wire pretension and differential tension can dominate the structural behavior, leading to shaft deflection and potential buckling. Therefore, design choices related to disk geometry, wire routing, hole dimensions, and shaft thickness must be made with manufacturability and stability explicitly considered. Basic Design of the Robotic Instrument: 2-DOF Bending and Shaft Rotation Axis with a Microgripper (OD 0.9 mm) The baseline robotic instrument design integrates a 2-DOF disk-stacked bending mechanism, a shaft rotation axis, and a microgripper within an OD of 0.9 mm. Two approaches can be considered for implementing a shaft rotation axis: rotating the entire instrument unit, or rotating only the shaft and the distal section. In the approach where the entire instrument unit is rotated, the robotic device—including both the instrument unit and the motor unit—inevitably becomes larger. When such a robotic device—comprising the instrument unit and motor unit mounted at the distal end of a positioning arm—is enlarged, there is concern that the increased size and mass may degrade rotational axis accuracy and lead to increased deflection of the positioning arm due to higher payload, resulting in reduced positioning accuracy and a lower natural frequency of the system. Rotating the entire instrument unit is advantageous in terms of achieving unlimited or multiple revolutions. However, rotating only the shaft and the distal section is less suitable for multi-turn rotation. In leader–follower robotic systems, nevertheless, a shaft rotation range of approximately ±180° is generally sufficient for surgical manipulation. Therefore, we adopted a basic design in which only the shaft and the distal section are rotated, rather than rotating the entire instrument unit. During shaft rotation, torsional deformation of the wires passing through the shaft may cause relative sliding between wires; however, this effect can be absorbed by the inherent flexibility of the wires themselves and is not expected to constitute a critical issue within the intended rotation range and wire pretension levels. This design choice increases the design freedom of the positioning arm and enables a more compact follower robot. Figures 2–5 illustrate the overall system configuration. Figure 2 shows the detachable structure between the Instrument unit and the motor unit, and Fig. 3 presents the external view and a cross-sectional view of the Instrument unit. Since the drive pulley shafts (motor shafts) correspond to the orientation axes and the end-effector actuation axis, they are defined as 𝜃4 through 𝜃7, and their respective positive rotation directions are indicated in the same figure. Figure 4 shows an enlarged cross-sectional view of the Instrument unit, and Fig. 5 depicts the distal bending mechanism of the Instrument unit. The actuation scheme for the 2-DOF bending mechanism and the microgripper is based on the previously reported wire-driven approach [22]. Specifically, the forceps shaft is supported by a shaft clamping component that is rotationally supported by two bearings. The shaft is driven by a pair of spur gears (m=0.5, Z0=24, Z1=12). To secure approximately ±180∘ of shaft rotation, a speed-increasing mechanism with a ratio of 2 is employed. A key issue identified in our earlier multi-DOF drive studies is stick–slip behavior caused by circumferential friction at the wire entrance point when the shaft rotation is performed under a bent configuration [25]. In the present design, a bearing and a sleeve are introduced at the rotating part so that the drive wires can avoid circumferential sliding at the wire entrance point during shaft rotation, thereby mitigating stick–slip and undesired vibration. Four drive pulleys for the 2-DOF bending, the microgripper, and the shaft rotation are arranged at 90° intervals around the circumference. The drive motors are identical DC servo motors with an encoder and planetary gearbox (DC-Motor-DCX08M EBKL4.2V, PlanetGearbox-GPX08A1296:1, MagneticEncoder-ENX8MAG256IMP). The instrument unit is designed to be detachable from the motor unit using guide pins and a latch lever (see Fig. 2 (2)), which improves usability and maintainability. Miniaturization of an Ultra-Fine Robotic Instrument (OD 0.7 mm) Key Considerations and Challenges for Miniaturization To achieve miniaturization from OD 0.9 mm to OD 0.7 mm, we must address manufacturability, assembly feasibility, and mechanical performance simultaneously. (a) Disk geometry (C-type vs. V-type). For OD 0.9 mm, we used a cylindrical-surface contact disk-stacked bending mechanism (C-type), where the upper and lower faces of each disk are cylindrical surfaces oriented 90° apart, and bending is obtained by stacking disks with opposing cylindrical faces. While this structure enables smooth bending, it requires complex machining and the disk orientation is difficult to distinguish visually during assembly, increasing the risk of assembly errors. For OD 0.7 mm, we adopt a V-shaped surface–constrained disk-stacked bending mechanism (V-type), in which each disk has a planar surface and a V-shaped surface, and stacking is performed by opposing the planar and V-shaped surfaces while rotating the disk orientation by 90°. This geometry uses a simple single-side V-shaped feature, improves manufacturability, and dramatically enhances assembly efficiency because the machining surface and ridge direction are easily recognizable. In addition, compared with the cylindrical-surface design, the V-type disk geometry can reduce the shift of the instantaneous rotation center during bending, thereby decreasing the path-length difference between pull and release wires and reducing tension variation during bending. Although the theoretical shift can be close to zero in an ideal geometric model, a small finite curvature remains in practice due to manufacturing tolerances and edge rounding. Figure 6 shows the disk geometry of the V-type configuration. The disk has an outer diameter of 0.7 mm, a slanted surface angle of 15°, a maximum disk height of 0.15 mm, a wire-hole diameter of 0.17 mm, and a wire-hole arrangement diameter of 0.43 mm. For an outer diameter of 0.7 mm, securing a larger hole diameter is difficult due to spatial constraints; therefore, the wire-hole diameter was set to 0.17 mm. (b) Drive-wire hole diameter In the C-type configuration with an outer diameter of 0.9 mm, a wire-hole diameter of 0.20 mm was used in combination with a 0.15-mm-diameter stranded SUS304 wire (1×19). In contrast, for the V-type configuration with an outer diameter of 0.7 mm, the wire-hole diameter was reduced to 0.17 mm. In the C-type mechanism, the wire holes of adjacent disks remain directly opposed even in the bent configuration. However, in the V-type mechanism, although the wire holes are opposed in the initial straight configuration, misalignment occurs between adjacent disks as bending progresses. This misalignment effectively reduces the available hole diameter for wire passage. Therefore, for the V-type configuration, the maximum wire diameter that can pass through the disk stack during bending was investigated. Figure 7 illustrates the maximum wire diameter that can pass through the guide wire holes when the disks are tilted to the maximum angle of 7.5°. The calculations were performed according to the procedure described below and subsequently verified by geometric drawing. The disks are assumed to rotate about the ridge line of the V-shaped surface. The drive wire is assumed to deform into a circular arc with constant curvature while passing through the guide wire holes of the bending mechanism, and to exit perpendicular to the upper surface of the distal disk of the bending disk pair. In other words, the wire is assumed to contact the outer side of the curvature at the exit of the guide wire hole as a tangent. Under this condition, the limiting case for wire passage occurs when the wire becomes tangent to the inner edge of the guide wire hole exit on the upper surface of the proximal disk in the bending disk pair. This configuration defines the maximum wire diameter that can pass through the disks. Based on the disk height, maximum thickness, hole diameter, hole position, and tilt angle, the maximum allowable wire diameter can be calculated. The derived maximum wire diameter was also confirmed by geometric drawing, as shown in Fig. 7. The resulting maximum allowable wire diameters were approximately 0.158 mm on the pull side and 0.161 mm on the release side, indicating that a wire with a diameter of 0.15 mm remains potentially applicable. However, because the effects of component manufacturing tolerances, assembly accuracy, friction, and wear are not considered in this analysis, experimental verification using an actual prototype is required to confirm whether reliable assembly and bending motion can be achieved. (c) Shaft stiffness and buckling During surgical manipulation, external forces (lateral loads) are applied to the distal end of the shaft, resulting in elastic deflection. In addition, axial compressive loads are generated in the shaft due to the tension of the drive wires required to actuate the bending mechanism, raising concerns about potential buckling of the shaft. In previous implementations with an outer diameter of 0.9 mm, no noticeable issues related to shaft deflection or buckling were observed. However, as the shaft diameter is reduced, a decrease in shaft stiffness is unavoidable. Therefore, theoretical evaluations of shaft deflection and buckling were conducted at the design stage. Let the outer diameter and inner diameter of the shaft be denoted by D and d , respectively. The second moment of area I of the shaft cross section is given by: $$\:I=\frac{\pi\:}{64}\left({D}^{4}-{d}^{4}\right)$$ 1 Assuming a lateral tip load \(\:{F}_{tip}\) , a shaft length L , Young’s modulus E , and a cantilever beam condition with an effective length factor K = 2, the tip deflection δ of the shaft can be expressed as: $$\:\delta\:=\frac{F{L}^{3}}{48EI}$$ 2 The Euler buckling load \(\:{P}_{cr}\) is given by: $$\:{P}_{cr}=\frac{{\pi\:}^{2}EI}{{\left(2L\right)}^{2}}$$ 3 In the current design with an outer diameter of 0.9 mm, a stainless-steel tube with dimensions OD 0.9 mm / ID 0.76 mm (wall thickness 0.07 mm) was used as instrument shaft (see Fig. 2). For an outer diameter of 0.7 mm, two candidate shaft geometries were considered as implementable designs: OD 0.7 mm / ID 0.6 mm (wall thickness 0.05 mm) and OD 0.7 mm / ID 0.5 mm (wall thickness 0.10 mm). The shaft length was set to 30 mm. The shaft material is austenitic stainless steel AISI 304 (equivalent to JIS SUS304), and its Young’s modulus was assumed to be 193 GPa. In retinal microsurgery, surgical manipulation is generally performed with tool–tissue interaction forces on the order of a few millinewtons. Previous experimental studies on membrane peeling have reported that typical instrument–tissue interaction forces are below 10 mN [ 27 , 28 ]; therefore, 10 mN was selected as a conservative upper-bound load for structural evaluation. To ensure sufficient structural robustness of the device, the tip deflection was therefore evaluated under a conservative lateral tip load of \(\:{F}_{tip}\) = 10 mN. The calculated results are summarized in Table 1. The deflection ratio and buckling load ratio are normalized with respect to the shaft with an outer diameter of 0.9 mm. When normalized to the OD 0.9 mm shaft, the second moment of area is reduced to 0.343 for the OD 0.7 mm / ID 0.6 mm shaft and to 0.551 for the OD 0.7 mm / ID 0.5 mm shaft. Consequently, the tip deflection under the same bending load increases by approximately 2.9 times for the ID 0.6 mm shaft and by 1.8 times for the ID 0.5 mm shaft, while the corresponding Euler buckling loads decrease to approximately 1/2.9 and 1/1.8 of the baseline value, respectively. Considering that the diameter of retinal blood vessels is typically on the order of 50–100 µm, it is desirable for the shaft deflection to remain within a comparable or smaller range. Based on these considerations, the OD 0.7 mm / ID 0.6 mm shaft was deemed unsuitable, and the OD 0.7 mm / ID 0.5 mm shaft was selected for implementation. With regard to buckling, to avoid buckling induced by the compressive load resulting from the initial pretension of the four bending drive wires, the initial pretension per wire should be limited such that the resulting axial compressive load acting on the shaft remains at least below one quarter of the Euler buckling load. This criterion is based on a conservative approximation that the axial compressive load acting on the shaft is equal to the sum of the tensions of the four bending drive wires. In general, wire-driven joint mechanisms do not exhibit tension variation solely as a function of joint angle. When an external load torque is applied, the tension in the pull-side wires increases by + ΔT, while the tension in the opposing wires decreases by − ΔT. When the tension variation ΔT is less than or equal to the initial pretension, the maximum axial compressive load acting on the shaft is governed by the initial pretension. Accordingly, under this condition, the axial compressive load does not exceed the value determined by the initial wire pretension. In the case of the 0.9-mm-OD instrument employing a variable neutral-line mechanism (C-type disk-stacked bending mechanism), an increase in bending angle leads to an increase in wire path length [ 22 ], which can result in an increase in wire tension. Therefore, tension variation associated with instrument posture must be carefully considered. In contrast, for the 0.7-mm-OD instrument employing the V-type disk-stacked bending mechanism, as described above, changes in wire path length can be minimized, which is advantageous with respect to buckling stability. Based on the above discussion of buckling load, the initial wire pretension is appropriately set during the assembly process. Basic Design of the Ultra-Fine Robotic Instrument (OD 0.7 mm) Based on the key points and challenges associated with the above miniaturization, the basic design of the 0.7‑mm‑OD robotic instrument was developed. The fundamental configuration follows that of Figs. 2 through 4. Figure 8 shows the distal bending mechanism of the 0.7‑mm‑OD instrument unit corresponding to Fig. 5. The coordinate system defined for the bending mechanism is also shown. The design is extended to a 0.7 mm outer diameter by introducing V‑type disk elements and modifying the shaft‑clamp geometry. The overall architecture of the instrument unit and motor unit remains identical to that of the 0.9 mm system, preserving consistent modularity and reusability of the motor‑unit platform. The primary dimensional change is the reduction of the shaft‑clamping hole from 0.9 mm to 0.7 mm, accompanied by corresponding updates to the disk stack and wire‑routing features. To accommodate miniaturization, the microgripper was radially scaled down to fit within a diameter of 0.7 mm while approximately maintaining its overall length. This design demonstrates that an ultra‑fine robotic instrument providing distal 3‑DOF orientation and an end‑effector can be realized while maintaining manufacturability and robustness suitable for ophthalmic MIMS. Results and Discussion (Prototyping and Validation) Based on the design concepts and design results presented in the previous section, components with outer diameters of 0.9 mm and 0.7 mm were fabricated and assembled. In addition, basic functional performance of the mechanical system was evaluated to validate manufacturability, assemblability, and fundamental operational feasibility. Prototyping Component materials and fabrication methods Key aspects of component fabrication are described below. The C-type and V-type disk elements were fabricated by precision machining of austenitic stainless steel AISI 304, according to the geometries shown in Figs. 1 and 6. Figure 9 shows the fabricated disk elements. Although the ridge line of the V-type disk element is ideally sharp, a minimal fillet radius was introduced to remove machining burrs and to avoid deformation caused by external forces. As a result, the actual disk height is slightly smaller than the nominal design value. It can be confirmed that the geometry of the V-type disk is easier to identify than that of the C-type disk. In previous prototypes, the main components of the Instrument unit were fabricated using a 3D printer. However, in the present study, the main structural components and the shaft clamp components—whose dimensional accuracy and stiffness are critical for assembly accuracy and motion precision— were fabricated by machining ABS resin, which provides sufficient stiffness for prototyping while allowing rapid and precise fabrication. The remaining components were fabricated using a stereolithography 3D printer (Formlabs Form 3) with Rigid 4000 resin. For the drive wires, commercially available wires were selected to ensure practical availability. The candidates included a 0.15-mm-diameter stranded AISI 304 wire (1×19), a 0.125-mm-diameter Ni–Ti wire, and a 0.1-mm-diameter Ni–Ti wire. Initial pretension setting As discussed in the previous section, the setting of the initial wire pretension is critically important. However, because the component dimensions are extremely small, precise pretension adjustment requires special fixtures. In addition, the method used to fix the drive wires to the drive pulleys during pretensioning is also important, while direct measurement of the wire tension after fixation is practically difficult. Therefore, the initial pretension was applied using the following simple and reproducible procedure. The method for fixing the wire to the pulley is shown in Fig. 10. The wire is passed through two holes provided in the drive pulley, folded back, and a weight is attached to the proximal end of the wire to apply a constant tensile load via gravity. While maintaining this condition, the wire is bonded through a hole on the side surface of the drive pulley using a cyanoacrylate adhesive. Thus, the initial pretension can be adjusted by selecting the mass of the attached weight. However, due to wire bending at the wire entrance point and friction at the folded section around the drive pulley, the tensile force applied at the distal end does not directly correspond to the gravitational force of the attached weight. To account for this effect, an identical wire routing configuration to that shown in Fig. 9 was constructed, and the relationship between the applied weight and the resulting distal wire tension was experimentally measured. Specifically, the proximal weight required to lift a distal weight was evaluated. As a result, it was found that approximately 40% of the applied gravitational force was transmitted as tensile force for the 0.15-mm-diameter AISI 304 wire, and approximately 33% for the 0.1-mm-diameter Ni–Ti wire. Based on these results, an initial pretension was applied using a 0.15-mm-diameter AISI 304 wire with a 300-g weight for the 0.9-mm-OD instrument, and a 0.1-mm-diameter Ni–Ti wire with a 200-g weight for the 0.7-mm-OD instrument. Table 2 summarizes the resulting initial pretension and the corresponding safety factor with respect to the Euler buckling load. In both cases, the safety factor was 1.78. Although this safety factor may not be sufficiently large in a strict structural design sense, the drive wires pass through the interior of the shaft; therefore, when shaft deflection occurs, the wire tension is expected to act as a restoring force against lateral deformation. Accordingly, it is assumed that plastic deformation or wire fracture will not occur. The adequacy of this assumption will be further evaluated based on whether any adverse effects arise during manipulation and operation in subsequent experiments. Assembly procedure Assembly of the 2-DOF bending mechanism and the microgripper was performed for both the 0.9-mm-OD and 0.7-mm-OD instruments under a stereo microscope using tweezers, without any special assembly jigs. Assembly of the entire Instrument unit could also be completed using simple jigs fabricated as needed and standard tools. For the 2-DOF bending mechanism with an outer diameter of 0.9 mm, a 0.15-mm-diameter AISI 304 wire was used, while a 0.125-mm-diameter Ni–Ti wire was used for the microgripper. For the 2-DOF bending mechanism with an outer diameter of 0.7 mm, assembly was feasible using a 0.15-mm-diameter AISI 304 wire as well as 0.125-mm- and 0.1-mm-diameter Ni–Ti wires; however, considering the smoothness of the bending mechanism motion, a 0.1-mm-diameter Ni–Ti wire was selected in this study. A 0.1-mm-diameter Ni–Ti wire was also used for the microgripper of the 0.7-mm-OD instrument. Fixation of the drive wires for the 2-DOF bending mechanism to the drive pulleys was performed in accordance with the initial pretension setting method described above. The drive wires for the microgripper were fixed to the drive pulleys with zero initial pretension, with the gripper assembled in the open state. Prototyping of the robotic instruments (OD 0.9 mm and OD 0.7 mm) Based on the component selection and assembly procedures described above, the robotic instruments were assembled. Figure 11 shows the fabricated instrument unit with an outer diameter of 0.7 mm together with the motor unit, and Fig. 12 presents photographs of the bending mechanisms with outer diameters of 0.9 mm and 0.7 mm. Preliminary Evaluation of the Mechanical System and Discussion In this section, in order to assess the potential for future clinical application, and because the fundamental performance data of the 2-DOF bending mechanism with an outer diameter of 0.9 mm have already been obtained in previous studies [22－24], the evaluation focuses primarily on the robotic instrument with an outer diameter of 0.7 mm. The experimental results presented below are described using the drive axes (𝜃4–𝜃7) and the coordinate system of the robotic instrument shown in Figs. 3 and 8. Specifically, counterclockwise rotation of the drive pulleys (motor shafts) is defined as the positive direction; the shaft rotation axis is defined as rotation about the z-axis (roll); and the two bending degrees of freedom are defined as rotations about the x-axis (pitch) and the y-axis (yaw). Motion results were recorded using a camera placed in the positive y-direction, capturing either video sequences or still images. Furthermore, still images, or still frames extracted from video recordings, were used to obtain dimensional data by referencing scale markings visible in the images and measuring them using 2D CAD software. Precise measurements with an accuracy of 0.001mm are difficult due to optical resolution limits and perspective effects. Therefore the measurement results are reported with a resolution of 0.01 mm. Basic operational evaluation of bending, rotation, and grasping functions Each axis was actuated using motor drive to perform shaft rotation, two-axis bending, and grasping motions, and it was confirmed that all functions operated smoothly (see Additional file 1). Figure 13 shows photographs of the instrument during bending actuation in the pitch and yaw directions, confirming the effectiveness of the bending function. Figure 14 shows the relationship between the drive pulley rotation angle and the resulting bending angle. As the drive pulley rotation angle increases, the bending angle increases approximately linearly; however, hysteresis of approximately ±5° is observed, which is presumed to be caused by friction between the wires and the disk holes. This level of hysteresis is considered acceptable for preliminary validation; however, further reduction through surface finishing, chamfering or lubrication will be investigated in future work. Figure 15 shows photographs confirming the grasping function. Although the opening angle is relatively small, approximately 10°, it is confirmed that the grasping function operates effectively. In the present evaluation of the microgripper function, quantitative measurement of the grasping force was beyond the scope of this study. Instead, the evaluation focused on deformation in the grasped state and the confirmation of stable grasping behavior. Quantitative evaluation of grasping force will be addressed in future work. Evaluation of shaft deflection and whirling during operation As analyzed in the previous section, shaft deflection due to buckling induced by wire tension is a potential concern associated with miniaturization. The following describes the evaluation results of shaft deflection during each type of operation. Shaft deflection during bending operation was measured from Fig. 14, and shaft deflection during grasping operation was measured from Fig. 15. For shaft whirling during rotation, videos were recorded while rotating the shaft from 0° to ±160° at a rotational speed of 10°/s under two conditions: a straight configuration with a bending angle of 0°, and a bent configuration with −45° pitch bending (about the x-axis). From these videos, the states exhibiting the maximum whirling displacement of the shaft were extracted. Figure 16 shows the maximum whirling condition observed during shaft rotation. Table 3 summarizes the shaft deflection and whirling displacement during bending, grasping, and shaft rotation operations. While the deflection during bending and grasping operations remains below approximately 0.06 mm, it is confirmed that shaft whirling with a maximum displacement amplitude of 0.1–0.23 mm occurs during shaft rotation. Comparing the straight and bent configurations, the total whirling amplitude is smaller in the bent configuration. Possible causes of the observed whirling include changes in wire tension associated with shaft rotation and bending actuation, as well as fabrication and assembly tolerances of components related to the shaft rotation axis. Identifying the specific causes requires measurements under various conditions, including three-dimensional measurements rather than single-direction observations and evaluation under different initial pretension settings; therefore, this issue is considered a subject for future work. From the perspective of future clinical application, shaft whirling of 0.1–0.23 mm during shaft rotation may have a limited impact on surgical procedures that do not involve changes in instrument orientation; however, when orientation changes are involved, the effects on operability and task performance must be carefully evaluated. In human-in-the-loop scenarios, such whirling is not expected to constitute a critical issue, as the surgeon can naturally compensate for the resulting deviations. Furthermore, for autonomous or semi-autonomous targeting, real-time compensation may be feasible, provided that the elastic deformation behavior is sufficiently consistent and predictable. Based on the above basic operational evaluation of the bending, rotation, and grasping functions, it was confirmed that the newly designed 3-DOF robotic instrument with a microgripper and an outer diameter of 0.7 mm has no critical issues from a mechanical design and prototyping perspective. Although the present results demonstrate the fundamental feasibility of the proposed ultra-fine robotic instrument, the evaluation was conducted under limited experimental conditions and is based on N = 1 measurements. Further systematic evaluation under diverse conditions and with larger sample sizes is required to fully characterize the performance and robustness of the system. Conclusion This paper presented the design and miniaturization of an ultra-fine multi-DOF robotic instrument for ophthalmic MIMS. The proposed instrument integrates a 2-DOF bending mechanism, a shaft rotation axis, and a microgripper within an outer diameter of 0.7 mm by introducing a surface-constrained (V-type) disk-stacked bending mechanism. Design considerations for miniaturization were addressed through (i) improved manufacturability and assembly efficiency of the disk elements, (ii) evaluation of wire passability through reduced guide-hole diameters, and (iii) quantitative assessment of shaft stiffness and Euler buckling under wire pretension and a conservative lateral tip load of 10 mN. Full-scale prototypes of both 0.9-mm-OD and 0.7-mm-OD instruments were fabricated and assembled, and basic bending, rotation, and grasping functions were experimentally verified. The 0.7-mm-OD instrument exhibited smooth pitch/yaw bending and grasping motion with a bending hysteresis of approximately ± 5°. Shaft deflection during bending and grasping remained below approximately 0.06 mm, while whirling during shaft rotation reached displacement amplitudes of 0.1–0.23 mm, indicating that further refinement of shaft rotation components and tension management will be beneficial for highly precise orientation control. Overall, the results demonstrate the feasibility of an ultra-fine robotic instrument with distal 3-DOF orientation and an end-effector suitable for ophthalmic MIMS, and clarify key design trade-offs among miniaturization, stiffness, wire routing, pretension, and stability. The insights obtained in this study provide a foundation for future extensions, including more comprehensive experimental validation, optimization of rotational stability, and integration into teleoperated ophthalmic robotic systems. Abbreviations Minimally invasive surgery MIS Minimally invasive microsurgery MIMS Degree-of-freedom DOF Remote center of motion RCM Declarations Author contributions M.J. led the conceptual planning and mechanical design of the robotic instrument, planned and conducted the verification experiments, performed the data analysis, and drafted the initial manuscript. I.I. played a key role in defining the specifications and requirements for the conceptual design. R.N. contributed to the development of the control system, including the design and implementation of the control software. All authors revised the manuscript and reviewed and approved the final version. Acknowledgments The authors acknowledge the support of the Japan Keirin Autorace Foundation (JKA). Competing interests The authors declare that they have no competing interests. Availability of data and materials Not applicable. Ethics approval and consent to participate Not applicable. Funding This research was supported by the subsidy program (individual research) of the Japan Keirin Autorace Foundation (JKA). References Intuitive Surgical, Inc., “da Vinci Surgical System.” Available: https://www.intuitive.com. Accessed: Jan. 13, 2026. M. S. Khan, M. R. Elhage, A. Challacombe, et al., “Technical review of the da Vinci surgical telemanipulator,” Int. J. Med. Robot. Comput. Assist. Surg., vol. 8, no. 4, pp. 467–476, 2012, doi:10.1002/rcs.1468. P. Probst, J. Keller, L. Müller, et al., “A review of the role of robotics in surgery: to da Vinci and beyond,” Frontiers in Surgery, vol. 10, Art. no. 1198574, 2023, doi:10.3389/fsurg.2023.1198574. Microsure B.V., “MUSA-2, MUSA-3.” Available: https://microsure.nl/. Accessed: Jan. 13, 2026. 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Res., vol. 6, nos. 1–2, 2140001, 2021. M. Jinno et al., “An integrated high-dexterity cooperative robotic assistant for intraocular micromanipulation,” in Proc. IEEE ICRA, 2021. M. Jinno and I. Iordachita, “Microgripper using flexible wire hinge for robotic intraocular snake,” in Proc. IEEE ICRA, 2022. M. Jinno, R. Nonoyama, and I. Iordachita, “Multi-degree-of-freedom bending mechanisms and drive mechanisms for robot-assisted minimally invasive microsurgery,” in Proc. IEEE/ASME AIM, 2025. J. Song, C. Gonenc, J. Guo, I. Iordachita, “Intraocular Snake Integrated with the Steady-Hand Eye Robot for Assisted Retinal Microsurgery,” in Proc. IEEE ICRA, 2017. S. Sunshine et al., “A force-sensing microsurgical instrument that detects forces below human tactile sensation,” Retina, vol. 33, no. 1, pp. 200–206, 2013, doi:10.1097/IAE.0b013e3182625d2b. P. K. Gupta, P. S. Jensen, and E. de Juan Jr., “Surgical forces and tactile perception during retinal microsurgery,” in Proc. MICCAI, LNCS vol. 1679, Springer, 1999. Tables Tables 1 to 3 are available in the Supplementary Files section. Additional Declarations No competing interests reported. Supplementary Files Table1.pptx Table 1. Comparison of bending stiffness, tip deflection, and buckling load Table2.pptx Table 2 Calculated safety factors against buckling due to the initial wire pretension Table3.pptx Table 3 Maximum shaft deflection and whirling displacement characteristics Additionalfile1.mp4 Cite Share Download PDF Status: Under Review Version 1 posted Editorial decision: Revision requested 31 Mar, 2026 Reviews received at journal 31 Mar, 2026 Reviewers agreed at journal 12 Mar, 2026 Reviews received at journal 09 Mar, 2026 Reviewers agreed at journal 09 Feb, 2026 Reviewers invited by journal 07 Feb, 2026 Editor assigned by journal 07 Feb, 2026 Submission checks completed at journal 06 Feb, 2026 First submitted to journal 03 Feb, 2026 You are reading this latest preprint version Research Square lets you share your work early, gain feedback from the community, and start making changes to your manuscript prior to peer review in a journal. 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Also discoverable on Platform About Our Team In Review Editorial Policies Advisory Board Help Center Resources Author Services Accessibility API Access RSS feed Manage Cookie Preferences © Research Square 2026 | ISSN 2693-5015 (online) Privacy Policy Terms of Service Do Not Sell My Personal Information {"props":{"pageProps":{"initialData":{"identity":"rs-8771621","acceptedTermsAndConditions":true,"allowDirectSubmit":false,"archivedVersions":[],"articleType":"Research Article","associatedPublications":[],"authors":[{"id":588315502,"identity":"cf71f303-65e3-4ca3-80b2-b9e87f621c61","order_by":0,"name":"Makoto 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07:43:41","extension":"pdf","order_by":0,"title":"","display":"","copyAsset":false,"role":"manuscript-pdf","size":8946600,"visible":true,"origin":"","legend":"","description":"","filename":"manuscript.pdf","url":"https://assets-eu.researchsquare.com/files/rs-8771621/v1/e8baaff9-6b1f-4244-8d00-c7d6efe70853.pdf"},{"id":102747069,"identity":"3bcd4c76-8dd7-4c49-9325-99f5b6702cb4","added_by":"auto","created_at":"2026-02-16 09:03:43","extension":"pptx","order_by":1,"title":"","display":"","copyAsset":false,"role":"supplement","size":39715,"visible":true,"origin":"","legend":"\u003cp\u003e\u003cstrong\u003eTable 1. Comparison of bending stiffness, tip deflection, and buckling load\u003c/strong\u003e\u003c/p\u003e","description":"","filename":"Table1.pptx","url":"https://assets-eu.researchsquare.com/files/rs-8771621/v1/2f5f4174dc3c904ba384ee53.pptx"},{"id":102593531,"identity":"8019d5db-d1b1-44be-8dd9-029f1234c738","added_by":"auto","created_at":"2026-02-13 11:50:37","extension":"pptx","order_by":2,"title":"","display":"","copyAsset":false,"role":"supplement","size":40755,"visible":true,"origin":"","legend":"\u003cp\u003e\u003cstrong\u003eTable 2 Calculated safety factors against buckling due to the initial wire pretension\u003c/strong\u003e\u003c/p\u003e","description":"","filename":"Table2.pptx","url":"https://assets-eu.researchsquare.com/files/rs-8771621/v1/4076fd92dca422bee380d1ba.pptx"},{"id":102593533,"identity":"efbf21fb-53e5-4516-bbfb-2040538af9f7","added_by":"auto","created_at":"2026-02-13 11:50:37","extension":"pptx","order_by":3,"title":"","display":"","copyAsset":false,"role":"supplement","size":40685,"visible":true,"origin":"","legend":"\u003cp\u003e\u003cstrong\u003eTable 3 Maximum shaft deflection and whirling displacement characteristics\u003c/strong\u003e\u003c/p\u003e","description":"","filename":"Table3.pptx","url":"https://assets-eu.researchsquare.com/files/rs-8771621/v1/2524d8b8c16aa2c876f66150.pptx"},{"id":102593548,"identity":"213651fc-9e2e-41e2-a221-ebc56876c76c","added_by":"auto","created_at":"2026-02-13 11:50:39","extension":"mp4","order_by":4,"title":"","display":"","copyAsset":false,"role":"supplement","size":44912048,"visible":true,"origin":"","legend":"","description":"","filename":"Additionalfile1.mp4","url":"https://assets-eu.researchsquare.com/files/rs-8771621/v1/287abbe9ea9084132ee19184.mp4"}],"financialInterests":"No competing interests reported.","formattedTitle":"Design and Miniaturization of an Ultra-Fine Multi-Degree-of-Freedom Robotic Instrument for Ophthalmic Minimally Invasive Microsurgery","fulltext":[{"header":"Introduction","content":"\u003cp\u003eMIS has fundamentally reshaped modern surgical practice by reducing patient trauma and enhancing postoperative quality of life. These benefits have been further amplified in laparoscopic surgery through the widespread clinical adoption of teleoperated robotic systems, most notably the da Vinci Surgical System, which offers superior dexterity, motion scaling, and ergonomic control within confined operative spaces [1,2]. As surgical robotics continues to evolve, there is an increasing expectation that its application will extend beyond conventional MIS toward microsurgical domains, where the demands for precision, stability, and dexterity often exceed the capabilities of manual techniques [3].\u003c/p\u003e\n\u003cp\u003eDespite this progress, the clinical translation of robotic assistance to microsurgery remains uneven and limited in routine practice. Although several robotic systems have been effectively implemented for microsurgical tasks in open environments [4,5], minimally invasive robotic systems designed to function within highly constrained spaces have not yet achieved widespread clinical adoption. Within this landscape, teleoperated leader\u0026ndash;follower robotic architectures offer a compelling solution for robot-assisted MIMS, as they enable precise motion scaling and tremor suppression while preserving intuitive surgeon control.\u003c/p\u003e\n\u003cp\u003eAmong microsurgical procedures, vitreoretinal surgery represents one of the most demanding applications in terms of invasiveness and precision. Accordingly, a wide range of robotic systems has been developed to assist retinal surgery [6], with feasibility demonstrated through in vivo animal experiments [7\u0026ndash;9] as well as first-in-human clinical studies [10,11]. However, many of these systems rely on straight, rigid instruments, which inherently restrict achievable approach angles and limit dexterity within the confined intraocular space.\u003c/p\u003e\n\u003cp\u003eTo address these limitations, snake-like and continuum robotic instruments have attracted increasing attention for retinal surgery. By providing distal bending capability, such instruments enable more favorable approach angles to anterior retinal regions and facilitate technically challenging procedures, including retinal vein cannulation and membrane peeling. Consequently, several bending-type robotic instruments, including snake-like and continuum designs, have been proposed for vitreoretinal applications [12\u0026ndash;14], underscoring the potential of articulated and continuously deformable distal mechanisms for enhancing dexterity within the confined intraocular workspace.\u003c/p\u003e\n\u003cp\u003eJoint mechanisms for robotic surgical instruments can be broadly classified into hinge-based articulated mechanisms and continuum bending mechanisms that undergo continuous deformation into an arc shape. For laparoscopic surgical instruments with outer diameters of approximately 7 mm, hinge-based joint mechanisms are commonly employed. However, when accounting for the strength requirements at hinge joints and the associated fabrication complexity, realizing submillimeter-scale instruments with outer diameters below 1 mm using hinge-based designs becomes technically impractical.\u003c/p\u003e\n\u003cp\u003eFor MIMS, bending-type mechanisms with simpler structural configurations and component geometries are generally more suitable for miniaturization and diameter reduction. Bending-type mechanisms can be further categorized into three principal approaches. The first is the notched-tube type, wherein a tubular structure is locally notched to create flexible hinge regions and actuated by tendons or wires [13\u0026ndash;15]. The second is the concentric-tube type, where multiple pre-curved, flexible concentric tubes are bent through relative rotation and translation [12,16\u0026ndash;17]. The third is the stacked-element type, in which discrete elements such as rings, vertebrae, or disks are stacked and bent via tendon or wire actuation [18\u0026ndash;21].\u003c/p\u003e\n\u003cp\u003eAlthough notched-tube-type mechanisms have been realized with outer diameters on the order of 1 mm, their clinical applicability remains extremely limited, and further diameter reduction is considered challenging due to fabrication constraints and insufficient structural strength. Concentric-tube-type mechanisms have achieved submillimeter-scale outer diameters; however, their large intrinsic curvature and structural characteristics make precise orientation control of end-effectors difficult within highly confined spaces. Consequently, achieving both extreme miniaturization and compact, controllable bending behavior remains challenging for both approaches.\u003c/p\u003e\n\u003cp\u003eIn contrast, stacked-element-type mechanisms are considered more amenable to miniaturization while preserving manufacturability, assembly feasibility, and bending performance through appropriate geometric design of the constituent elements. Based on the principle of the variable neutral-line mechanism [21], we therefore investigate a disk-stacked bending mechanism as an instrument unit for robot-assisted MIMS. In our previous studies, we developed an improved robotic intraocular snake and analyzed the kinematics and drive mechanisms of its highly dexterous distal unit, demonstrating the feasibility of an ultra-fine, disk-stacked 2-DOF bending mechanism with an outer diameter of 0.9 mm suitable for intraocular manipulation [22,23] (see Fig. 1(1) and (2)). We subsequently developed a microgripper compatible with this ultra-fine bending mechanism and demonstrated stable grasping capability even within an extremely constrained workspace [24] (see Fig. 1(3)). These studies confirmed that ultra-miniaturized bending mechanisms and distal end-effectors can be successfully implemented for intraocular robotic applications.\u003c/p\u003e\n\u003cp\u003eIn addition to the 2-DOF bending mechanism, and with future clinical applications in mind, we also proposed multi-DOF bending mechanisms and corresponding drive systems designed for integration with follower arms of teleoperated surgical robots [25]. However, for practical intraocular surgical applications, outer diameters of 0.9 mm inherently limit the range of feasible procedures, making further miniaturization necessary. Such diameter reduction is expected to introduce severe trade-offs among bending capability, shaft stiffness, wire routing, pretension, and buckling stability.\u003c/p\u003e\n\u003cp\u003eTo address these challenges, this study aims to present the design and miniaturization of an ultra-fine multi-DOF robotic instrument specifically intended for ophthalmic MIMS. The proposed instrument integrates a 2-DOF bending mechanism, a shaft rotation axis, and a microgripper within an outer diameter of 0.7 mm. To enhance manufacturability and assembly feasibility at this scale, a novel disk-stacked bending mechanism based on a surface-constrained disk geometry is introduced. In addition, the influences of shaft dimensions, wire routing configuration, and initial wire pretension on bending performance, distal tip deflection, and buckling behavior are systematically investigated through experiments using full-scale prototypes. Through these studies, this paper elucidates the key design trade-offs associated with ultra-fine robotic instruments and provides experimentally validated and practically applicable design guidelines for future ophthalmic MIMS systems.\u003c/p\u003e"},{"header":"Methods (Mechanical Design)","content":"\u003cp\u003eThis section discusses the design and miniaturization of 3-DOF robotic instruments with a microgripper. Here, \u0026ldquo;3-DOF\u0026rdquo; refers exclusively to distal orientation (pitch, yaw, roll) and does not include insertion/translation, remote center of motion (RCM), or end-effector actuation. Robotic assistance in surgery can be broadly categorized into (i) handheld devices, (ii) teleoperated (leader\u0026ndash;follower) robotic systems, and (iii) human\u0026ndash;robot cooperative control systems [26]. In laparoscopic surgery, teleoperated systems exemplified by the da Vinci platform have become the de facto standard in clinical practice due to their superior ergonomics, dexterity, and stable manipulation in the abdominal cavity. In minimally invasive microsurgery (MIMS), handheld, teleoperated, and cooperative control approaches are all conceivable; however, as clinical demands expand toward complex and prolonged tasks requiring high reproducibility, teleoperated systems are expected to become increasingly important. Moreover, considering future developments toward automation using AI-based perception and decision-making, the follower robot technology of teleoperated systems provides a natural basis for semiautonomous or supervised-autonomous microsurgical functions.\u003c/p\u003e\n\u003cp\u003eA typical follower robot for teleoperated MIMS consists of (1) a positioning arm providing an RCM constraint to maintain the insertion point (sclerotomy), and (2) a robotic instrument that provides distal orientation DOFs plus an end-effector. In this study, we focus on the design of the robotic instrument that integrates \u0026ldquo;3-DOF distal orientation + end-effector,\u0026rdquo; where the distal orientation is realized by a 2-DOF bending mechanism and a shaft rotation axis, and the end-effector is a microgripper.\u0026nbsp;\u003c/p\u003e\n\u003cp\u003e\u003cstrong\u003eDesign Requirements Derived from Ophthalmic MIMS\u003c/strong\u003e\u003c/p\u003e\n\u003cp\u003eOphthalmic MIMS, such as vitreoretinal surgery, imposes stringent constraints on instrument diameter, dexterity, and mechanical stability. First, the insertion diameter must be sufficiently small to be compatible with standard trocar-based workflows. While an outer diameter (OD) of 0.9 mm can be used in certain experimental or limited clinical scenarios, its applicability is restricted because it may require enlargement of the sclerotomy or reduce compatibility with standard tool exchanges. Reducing the OD to approximately 0.7 mm, corresponding to 22-gauge-class instrumentation, substantially broadens the range of potential ophthalmic procedures and facilitates integration into routine workflows. Second, the instrument must provide adequate distal dexterity to achieve favorable approach angles and stable manipulation inside the eyeball. For this purpose, we target a distal orientation structure that combines (i) 2-DOF bending (pitch and yaw) and (ii) a shaft rotation axis (roll) while maintaining an end-effector actuation DOF. Third, mechanical stability under wire-driven actuation is essential. In ultra-fine instruments, internal forces generated by wire pretension and differential tension can dominate the structural behavior, leading to shaft deflection and potential buckling. Therefore, design choices related to disk geometry, wire routing, hole dimensions, and shaft thickness must be made with manufacturability and stability explicitly considered.\u003c/p\u003e\n\u003cp\u003e\u003cstrong\u003eBasic Design of the Robotic Instrument: 2-DOF Bending and Shaft Rotation Axis with a Microgripper (OD 0.9 mm)\u003c/strong\u003e\u003c/p\u003e\n\u003cp\u003eThe baseline robotic instrument design integrates a 2-DOF disk-stacked bending mechanism, a shaft rotation axis, and a microgripper within an OD of 0.9 mm.\u0026nbsp;\u003c/p\u003e\n\u003cp\u003eTwo approaches can be considered for implementing a shaft rotation axis: rotating the entire instrument unit, or rotating only the shaft and the distal section. In the approach where the entire instrument unit is rotated, the robotic device\u0026mdash;including both the instrument unit and the motor unit\u0026mdash;inevitably becomes larger. When such a robotic device\u0026mdash;comprising the instrument unit and motor unit mounted at the distal end of a positioning arm\u0026mdash;is enlarged, there is concern that the increased size and mass may degrade rotational axis accuracy and lead to increased deflection of the positioning arm due to higher payload, resulting in reduced positioning accuracy and a lower natural frequency of the system. Rotating the entire instrument unit is advantageous in terms of achieving unlimited or multiple revolutions. However, rotating only the shaft and the distal section is less suitable for multi-turn rotation. In leader\u0026ndash;follower robotic systems, nevertheless, a shaft rotation range of approximately \u0026plusmn;180\u0026deg; is generally sufficient for surgical manipulation. Therefore, we adopted a basic design in which only the shaft and the distal section are rotated, rather than rotating the entire instrument unit. During shaft rotation, torsional deformation of the wires passing through the shaft may cause relative sliding between wires; however, this effect can be absorbed by the inherent flexibility of the wires themselves and is not expected to constitute a critical issue within the intended rotation range and wire pretension levels. This design choice increases the design freedom of the positioning arm and enables a more compact follower robot.\u003c/p\u003e\n\u003cp\u003eFigures 2\u0026ndash;5 illustrate the overall system configuration. Figure 2 shows the detachable structure between the Instrument unit and the motor unit, and Fig. 3 presents the external view and a cross-sectional view of the Instrument unit. Since the drive pulley shafts (motor shafts) correspond to the orientation axes and the end-effector actuation axis, they are defined as\u0026nbsp;𝜃4 through\u0026nbsp;𝜃7, and their respective positive rotation directions are indicated in the same figure. Figure 4 shows an enlarged cross-sectional view of the Instrument unit, and Fig. 5 depicts the distal bending mechanism of the Instrument unit. The actuation scheme for the 2-DOF bending mechanism and the microgripper is based on the previously reported wire-driven approach [22]. Specifically, the forceps shaft is supported by a shaft clamping component that is rotationally supported by two bearings. The shaft is driven by a pair of spur gears (m=0.5, Z0=24, Z1=12). To secure approximately \u0026plusmn;180∘\u0026nbsp;of shaft rotation, a speed-increasing mechanism with a ratio of 2 is employed. A key issue identified in our earlier multi-DOF drive studies is stick\u0026ndash;slip behavior caused by circumferential friction at the wire entrance point when the shaft rotation is performed under a bent configuration [25]. In the present design, a bearing and a sleeve are introduced at the rotating part so that the drive wires can avoid circumferential sliding at the wire entrance point during shaft rotation, thereby mitigating stick\u0026ndash;slip and undesired vibration. Four drive pulleys for the 2-DOF bending, the microgripper, and the shaft rotation are arranged at 90\u0026deg; intervals around the circumference. The drive motors are identical DC servo motors with an encoder and planetary gearbox (DC-Motor-DCX08M EBKL4.2V, PlanetGearbox-GPX08A1296:1, MagneticEncoder-ENX8MAG256IMP). The instrument unit is designed to be detachable from the motor unit using guide pins and a latch lever\u0026nbsp;(see Fig. 2 (2)), which improves usability and maintainability.\u003c/p\u003e\n\u003cp\u003e\u003cstrong\u003eMiniaturization of an Ultra-Fine Robotic Instrument (OD 0.7 mm)\u003c/strong\u003e\u003c/p\u003e\n\u003cp\u003e\u003cstrong\u003e\u003cem\u003eKey Considerations and Challenges for Miniaturization\u003c/em\u003e\u003c/strong\u003e\u003c/p\u003e\n\u003cp\u003eTo achieve miniaturization from OD 0.9 mm to OD 0.7 mm, we must address manufacturability, assembly feasibility, and mechanical performance simultaneously.\u003c/p\u003e\n\u003cp\u003e(a) Disk geometry (C-type vs. V-type).\u003cbr\u003e\u0026nbsp;For OD 0.9 mm, we used a cylindrical-surface contact disk-stacked bending mechanism (C-type), where the upper and lower faces of each disk are cylindrical surfaces oriented 90\u0026deg; apart, and bending is obtained by stacking disks with opposing cylindrical faces. While this structure enables smooth bending, it requires complex machining and the disk orientation is difficult to distinguish visually during assembly, increasing the risk of assembly errors. For OD 0.7 mm, we adopt a V-shaped surface\u0026ndash;constrained disk-stacked bending mechanism (V-type), in which each disk has a planar surface and a V-shaped surface, and stacking is performed by opposing the planar and V-shaped surfaces while rotating the disk orientation by 90\u0026deg;. This geometry uses a simple single-side V-shaped feature, improves manufacturability, and dramatically enhances assembly efficiency because the machining surface and ridge direction are easily recognizable. In addition, compared with the cylindrical-surface design, the V-type disk geometry can reduce the shift of the instantaneous rotation center during bending, thereby decreasing the path-length difference between pull and release wires and reducing tension variation during bending. Although the theoretical shift can be close to zero in an ideal geometric model, a small finite curvature remains in practice due to manufacturing tolerances and edge rounding.\u003c/p\u003e\n\u003cp\u003eFigure 6 shows the disk geometry of the V-type configuration. The disk has an outer diameter of 0.7 mm, a slanted surface angle of 15\u0026deg;, a maximum disk height of 0.15 mm, a wire-hole diameter of 0.17 mm, and a wire-hole arrangement diameter of 0.43 mm. For an outer diameter of 0.7 mm, securing a larger hole diameter is difficult due to spatial constraints; therefore, the wire-hole diameter was set to 0.17 mm.\u003c/p\u003e\n\u003cp\u003e(b) Drive-wire hole diameter\u003c/p\u003e\n\u003cp\u003eIn the C-type configuration with an outer diameter of 0.9 mm, a wire-hole diameter of 0.20 mm was used in combination with a 0.15-mm-diameter stranded SUS304 wire (1\u0026times;19). In contrast, for the V-type configuration with an outer diameter of 0.7 mm, the wire-hole diameter was reduced to 0.17 mm. In the C-type mechanism, the wire holes of adjacent disks remain directly opposed even in the bent configuration. However, in the V-type mechanism, although the wire holes are opposed in the initial straight configuration, misalignment occurs between adjacent disks as bending progresses. This misalignment effectively reduces the available hole diameter for wire passage.\u003c/p\u003e\n\u003cp\u003eTherefore, for the V-type configuration, the maximum wire diameter that can pass through the disk stack during bending was investigated.\u0026nbsp;\u003c/p\u003e\n\u003cp\u003eFigure 7 illustrates the maximum wire diameter that can pass through the guide wire holes when the disks are tilted to the maximum angle of 7.5\u0026deg;. The calculations were performed according to the procedure described below and subsequently verified by geometric drawing. The disks are assumed to rotate about the ridge line of the V-shaped surface. The drive wire is assumed to deform into a circular arc with constant curvature while passing through the guide wire holes of the bending mechanism, and to exit perpendicular to the upper surface of the distal disk of the bending disk pair. In other words, the wire is assumed to contact the outer side of the curvature at the exit of the guide wire hole as a tangent. Under this condition, the limiting case for wire passage occurs when the wire becomes tangent to the inner edge of the guide wire hole exit on the upper surface of the proximal disk in the bending disk pair. This configuration defines the maximum wire diameter that can pass through the disks.\u003c/p\u003e\n\u003cp\u003eBased on the disk height, maximum thickness, hole diameter, hole position, and tilt angle, the maximum allowable wire diameter can be calculated. The derived maximum wire diameter was also confirmed by geometric drawing, as shown in Fig. 7. The resulting maximum allowable wire diameters were approximately 0.158 mm on the pull side and 0.161 mm on the release side, indicating that a wire with a diameter of 0.15 mm remains potentially applicable. However, because the effects of component manufacturing tolerances, assembly accuracy, friction, and wear are not considered in this analysis, experimental verification using an actual prototype is required to confirm whether reliable assembly and bending motion can be achieved.\u003c/p\u003e\n\u003cp\u003e(c) Shaft stiffness and buckling\u003cbr\u003eDuring surgical manipulation, external forces (lateral loads) are applied to the distal end of the shaft, resulting in elastic deflection. In addition, axial compressive loads are generated in the shaft due to the tension of the drive wires required to actuate the bending mechanism, raising concerns about potential buckling of the shaft. In previous implementations with an outer diameter of 0.9 mm, no noticeable issues related to shaft deflection or buckling were observed. However, as the shaft diameter is reduced, a decrease in shaft stiffness is unavoidable. Therefore, theoretical evaluations of shaft deflection and buckling were conducted at the design stage. Let the outer diameter and inner diameter of the shaft be denoted by \u003cem\u003eD\u003c/em\u003e and \u003cem\u003ed\u003c/em\u003e, respectively. The second moment of area \u003cem\u003eI\u003c/em\u003e of the shaft cross section is given by:\u003c/p\u003e\n\u003cdiv id=\"Sec5\" class=\"Section2\"\u003e\n \u003cdiv id=\"Equ1\" class=\"Equation\"\u003e\n \u003cdiv class=\"mathdisplay\" id=\"FileID_Equ1\" name=\"EquationSource\"\u003e$$\\:I=\\frac{\\pi\\:}{64}\\left({D}^{4}-{d}^{4}\\right)$$\u003c/div\u003e\n \u003cdiv class=\"EquationNumber\"\u003e1\u003c/div\u003e\n \u003c/div\u003e\n \u003cp\u003eAssuming a lateral tip load \u003cspan class=\"InlineEquation\"\u003e\u003cspan class=\"mathinline\"\u003e\\(\\:{F}_{tip}\\)\u003c/span\u003e\u003c/span\u003e, a shaft length \u003cem\u003eL\u003c/em\u003e, Young\u0026rsquo;s modulus \u003cem\u003eE\u003c/em\u003e, and a cantilever beam condition with an effective length factor \u003cem\u003eK\u003c/em\u003e\u0026thinsp;=\u0026thinsp;2, the tip deflection \u003cem\u003e\u0026delta;\u003c/em\u003e of the shaft can be expressed as:\u003c/p\u003e\n \u003cdiv id=\"Equ2\" class=\"Equation\"\u003e\n \u003cdiv class=\"mathdisplay\" id=\"FileID_Equ2\" name=\"EquationSource\"\u003e$$\\:\\delta\\:=\\frac{F{L}^{3}}{48EI}$$\u003c/div\u003e\n \u003cdiv class=\"EquationNumber\"\u003e2\u003c/div\u003e\n \u003c/div\u003e\n \u003cp\u003eThe Euler buckling load \u003cspan class=\"InlineEquation\"\u003e\u003cspan class=\"mathinline\"\u003e\\(\\:{P}_{cr}\\)\u003c/span\u003e\u003c/span\u003e is given by:\u003c/p\u003e\n \u003cdiv id=\"Equ3\" class=\"Equation\"\u003e\n \u003cdiv class=\"mathdisplay\" id=\"FileID_Equ3\" name=\"EquationSource\"\u003e$$\\:{P}_{cr}=\\frac{{\\pi\\:}^{2}EI}{{\\left(2L\\right)}^{2}}$$\u003c/div\u003e\n \u003cdiv class=\"EquationNumber\"\u003e3\u003c/div\u003e\n \u003c/div\u003e\n \u003cp\u003eIn the current design with an outer diameter of 0.9 mm, a stainless-steel tube with dimensions OD 0.9 mm / ID 0.76 mm (wall thickness 0.07 mm) was used as instrument shaft (see Fig. 2). For an outer diameter of 0.7 mm, two candidate shaft geometries were considered as implementable designs: OD 0.7 mm / ID 0.6 mm (wall thickness 0.05 mm) and OD 0.7 mm / ID 0.5 mm (wall thickness 0.10 mm). The shaft length was set to 30 mm. The shaft material is austenitic stainless steel AISI 304 (equivalent to JIS SUS304), and its Young\u0026rsquo;s modulus was assumed to be 193 GPa.\u003c/p\u003e\n \u003cp\u003eIn retinal microsurgery, surgical manipulation is generally performed with tool\u0026ndash;tissue interaction forces on the order of a few millinewtons. Previous experimental studies on membrane peeling have reported that typical instrument\u0026ndash;tissue interaction forces are below 10 mN [\u003cspan class=\"CitationRef\"\u003e27\u003c/span\u003e, \u003cspan class=\"CitationRef\"\u003e28\u003c/span\u003e]; therefore, 10 mN was selected as a conservative upper-bound load for structural evaluation. To ensure sufficient structural robustness of the device, the tip deflection was therefore evaluated under a conservative lateral tip load of \u003cspan class=\"InlineEquation\"\u003e\u003cspan class=\"mathinline\"\u003e\\(\\:{F}_{tip}\\)\u003c/span\u003e\u003c/span\u003e = 10 mN. The calculated results are summarized in Table 1. The deflection ratio and buckling load ratio are normalized with respect to the shaft with an outer diameter of 0.9 mm. When normalized to the OD 0.9 mm shaft, the second moment of area is reduced to 0.343 for the OD 0.7 mm / ID 0.6 mm shaft and to 0.551 for the OD 0.7 mm / ID 0.5 mm shaft. Consequently, the tip deflection under the same bending load increases by approximately 2.9 times for the ID 0.6 mm shaft and by 1.8 times for the ID 0.5 mm shaft, while the corresponding Euler buckling loads decrease to approximately 1/2.9 and 1/1.8 of the baseline value, respectively.\u003c/p\u003e\n \u003cp\u003eConsidering that the diameter of retinal blood vessels is typically on the order of 50\u0026ndash;100 \u0026micro;m, it is desirable for the shaft deflection to remain within a comparable or smaller range. Based on these considerations, the OD 0.7 mm / ID 0.6 mm shaft was deemed unsuitable, and the OD 0.7 mm / ID 0.5 mm shaft was selected for implementation. With regard to buckling, to avoid buckling induced by the compressive load resulting from the initial pretension of the four bending drive wires, the initial pretension per wire should be limited such that the resulting axial compressive load acting on the shaft remains at least below one quarter of the Euler buckling load. This criterion is based on a conservative approximation that the axial compressive load acting on the shaft is equal to the sum of the tensions of the four bending drive wires. In general, wire-driven joint mechanisms do not exhibit tension variation solely as a function of joint angle. When an external load torque is applied, the tension in the pull-side wires increases by\u0026thinsp;+\u0026thinsp;\u0026Delta;T, while the tension in the opposing wires decreases by\u0026thinsp;\u0026minus;\u0026thinsp;\u0026Delta;T. When the tension variation \u0026Delta;T is less than or equal to the initial pretension, the maximum axial compressive load acting on the shaft is governed by the initial pretension. Accordingly, under this condition, the axial compressive load does not exceed the value determined by the initial wire pretension. In the case of the 0.9-mm-OD instrument employing a variable neutral-line mechanism (C-type disk-stacked bending mechanism), an increase in bending angle leads to an increase in wire path length [\u003cspan class=\"CitationRef\"\u003e22\u003c/span\u003e], which can result in an increase in wire tension. Therefore, tension variation associated with instrument posture must be carefully considered. In contrast, for the 0.7-mm-OD instrument employing the V-type disk-stacked bending mechanism, as described above, changes in wire path length can be minimized, which is advantageous with respect to buckling stability. Based on the above discussion of buckling load, the initial wire pretension is appropriately set during the assembly process.\u003c/p\u003e\n\u003c/div\u003e\n\u003ch3\u003eBasic Design of the Ultra-Fine Robotic Instrument (OD 0.7 mm)\u003c/h3\u003e\n\u003cp\u003eBased on the key points and challenges associated with the above miniaturization, the basic design of the 0.7‑mm‑OD robotic instrument was developed. The fundamental configuration follows that of Figs.\u0026nbsp;2 through 4. Figure\u0026nbsp;8 shows the distal bending mechanism of the 0.7‑mm‑OD instrument unit corresponding to Fig.\u0026nbsp;5. The coordinate system defined for the bending mechanism is also shown. The design is extended to a 0.7 mm outer diameter by introducing V‑type disk elements and modifying the shaft‑clamp geometry. The overall architecture of the instrument unit and motor unit remains identical to that of the 0.9 mm system, preserving consistent modularity and reusability of the motor‑unit platform. The primary dimensional change is the reduction of the shaft‑clamping hole from 0.9 mm to 0.7 mm, accompanied by corresponding updates to the disk stack and wire‑routing features. To accommodate miniaturization, the microgripper was radially scaled down to fit within a diameter of 0.7 mm while approximately maintaining its overall length.\u003c/p\u003e\n\u003cp\u003eThis design demonstrates that an ultra‑fine robotic instrument providing distal 3‑DOF orientation and an end‑effector can be realized while maintaining manufacturability and robustness suitable for ophthalmic MIMS.\u003c/p\u003e"},{"header":"Results and Discussion (Prototyping and Validation)","content":"\u003cp\u003eBased on the design concepts and design results presented in the previous section, components with outer diameters of 0.9 mm and 0.7 mm were fabricated and assembled. In addition, basic functional performance of the mechanical system was evaluated to validate manufacturability, assemblability, and fundamental operational feasibility.\u003c/p\u003e\n\u003cp\u003e\u003cstrong\u003ePrototyping\u003c/strong\u003e\u003c/p\u003e\n\u003cp\u003e\u003cstrong\u003e\u003cem\u003eComponent materials and fabrication methods\u003c/em\u003e\u003c/strong\u003e\u003c/p\u003e\n\u003cp\u003eKey aspects of component fabrication are described below.\u003c/p\u003e\n\u003cp\u003eThe C-type and V-type disk elements were fabricated by precision machining of austenitic stainless steel AISI 304, according to the geometries shown in Figs. 1 and 6. Figure 9 shows the fabricated disk elements. Although the ridge line of the V-type disk element is ideally sharp, a minimal fillet radius was introduced to remove machining burrs and to avoid deformation caused by external forces. As a result, the actual disk height is slightly smaller than the nominal design value. It can be confirmed that the geometry of the V-type disk is easier to identify than that of the C-type disk.\u003c/p\u003e\n\u003cp\u003eIn previous prototypes, the main components of the Instrument unit were fabricated using a 3D printer. However, in the present study, the main structural components and the shaft clamp components\u0026mdash;whose dimensional accuracy and stiffness are critical for assembly accuracy and motion precision\u0026mdash;\u0026nbsp;were fabricated by machining ABS resin, which provides sufficient stiffness for prototyping while allowing rapid and precise fabrication. The remaining components were fabricated using a stereolithography 3D printer (Formlabs Form 3) with Rigid 4000 resin.\u003c/p\u003e\n\u003cp\u003eFor the drive wires, commercially available wires were selected to ensure practical availability. The candidates included a 0.15-mm-diameter stranded AISI 304 wire (1\u0026times;19), a 0.125-mm-diameter Ni\u0026ndash;Ti wire, and a 0.1-mm-diameter Ni\u0026ndash;Ti wire.\u003c/p\u003e\n\u003cp\u003e\u003cstrong\u003e\u003cem\u003eInitial pretension setting\u003c/em\u003e\u003c/strong\u003e\u003c/p\u003e\n\u003cp\u003eAs discussed in the previous section, the setting of the initial wire pretension is critically important. However, because the component dimensions are extremely small, precise pretension adjustment requires special fixtures. In addition, the method used to fix the drive wires to the drive pulleys during pretensioning is also important, while direct measurement of the wire tension after fixation is practically difficult. Therefore, the initial pretension was applied using the following simple and reproducible procedure.\u003c/p\u003e\n\u003cp\u003eThe method for fixing the wire to the pulley is shown in Fig. 10. The wire is passed through two holes provided in the drive pulley, folded back, and a weight is attached to the proximal end of the wire to apply a constant tensile load via gravity. While maintaining this condition, the wire is bonded through a hole on the side surface of the drive pulley using a cyanoacrylate adhesive. Thus, the initial pretension can be adjusted by selecting the mass of the attached weight.\u003c/p\u003e\n\u003cp\u003eHowever, due to wire bending at the wire entrance point and friction at the folded section around the drive pulley, the tensile force applied at the distal end does not directly correspond to the gravitational force of the attached weight. To account for this effect, an identical wire routing configuration to that shown in Fig. 9 was constructed, and the relationship between the applied weight and the resulting distal wire tension was experimentally measured. Specifically, the proximal weight required to lift a distal weight was evaluated.\u003c/p\u003e\n\u003cp\u003eAs a result, it was found that approximately 40% of the applied gravitational force was transmitted as tensile force for the 0.15-mm-diameter AISI 304 wire, and approximately 33% for the 0.1-mm-diameter Ni\u0026ndash;Ti wire. Based on these results, an initial pretension was applied using a 0.15-mm-diameter AISI 304 wire with a 300-g weight for the 0.9-mm-OD instrument, and a 0.1-mm-diameter Ni\u0026ndash;Ti wire with a 200-g weight for the 0.7-mm-OD instrument.\u003c/p\u003e\n\u003cp\u003eTable 2 summarizes the resulting initial pretension and the corresponding safety factor with respect to the Euler buckling load. In both cases, the safety factor was 1.78. Although this safety factor may not be sufficiently large in a strict structural design sense, the drive wires pass through the interior of the shaft; therefore, when shaft deflection occurs, the wire tension is expected to act as a restoring force against lateral deformation. Accordingly, it is assumed that plastic deformation or wire fracture will not occur. The adequacy of this assumption will be further evaluated based on whether any adverse effects arise during manipulation and operation in subsequent experiments.\u003c/p\u003e\n\u003cp\u003e\u003cstrong\u003e\u003cem\u003eAssembly procedure\u003c/em\u003e\u003c/strong\u003e\u003c/p\u003e\n\u003cp\u003eAssembly of the 2-DOF bending mechanism and the microgripper was performed for both the 0.9-mm-OD and 0.7-mm-OD instruments under a stereo microscope using tweezers, without any special assembly jigs. Assembly of the entire Instrument unit could also be completed using simple jigs fabricated as needed and standard tools. For the 2-DOF bending mechanism with an outer diameter of 0.9 mm, a 0.15-mm-diameter AISI 304 wire was used, while a 0.125-mm-diameter Ni\u0026ndash;Ti wire was used for the microgripper. For the 2-DOF bending mechanism with an outer diameter of 0.7 mm, assembly was feasible using a 0.15-mm-diameter AISI 304 wire as well as 0.125-mm- and 0.1-mm-diameter Ni\u0026ndash;Ti wires; however, considering the smoothness of the bending mechanism motion, a 0.1-mm-diameter Ni\u0026ndash;Ti wire was selected in this study. A 0.1-mm-diameter Ni\u0026ndash;Ti wire was also used for the microgripper of the 0.7-mm-OD instrument. Fixation of the drive wires for the 2-DOF bending mechanism to the drive pulleys was performed in accordance with the initial pretension setting method described above. The drive wires for the microgripper were fixed to the drive pulleys with zero initial pretension, with the gripper assembled in the open state.\u003c/p\u003e\n\u003cp\u003e\u003cstrong\u003e\u003cem\u003ePrototyping of the robotic instruments (OD 0.9 mm and OD 0.7 mm)\u003c/em\u003e\u003c/strong\u003e\u003c/p\u003e\n\u003cp\u003eBased on the component selection and assembly procedures described above, the robotic instruments were assembled. Figure 11 shows the fabricated instrument unit with an outer diameter of 0.7 mm together with the motor unit, and Fig. 12 presents photographs of the bending mechanisms with outer diameters of 0.9 mm and 0.7 mm.\u003c/p\u003e\n\u003cp\u003e\u003cstrong\u003ePreliminary Evaluation of the Mechanical System and Discussion\u003c/strong\u003e\u003c/p\u003e\n\u003cp\u003eIn this section, in order to assess the potential for future clinical application, and because the fundamental performance data of the 2-DOF bending mechanism with an outer diameter of 0.9 mm have already been obtained in previous studies [22－24], the evaluation focuses primarily on the robotic instrument with an outer diameter of 0.7 mm.\u003c/p\u003e\n\u003cp\u003eThe experimental results presented below are described using the drive axes (𝜃4\u0026ndash;𝜃7) and the coordinate system of the robotic instrument shown in Figs. 3 and 8. Specifically, counterclockwise rotation of the drive pulleys (motor shafts) is defined as the positive direction; the shaft rotation axis is defined as rotation about the z-axis (roll); and the two bending degrees of freedom are defined as rotations about the x-axis (pitch) and the y-axis (yaw). Motion results were recorded using a camera placed in the positive y-direction, capturing either video sequences or still images. Furthermore, still images, or still frames extracted from video recordings, were used to obtain dimensional data by referencing scale markings visible in the images and measuring them using 2D CAD software. Precise measurements with an accuracy of 0.001mm are difficult due to optical resolution limits and perspective effects. Therefore the measurement results are reported with a resolution of 0.01 mm.\u003c/p\u003e\n\u003cp\u003e\u003cstrong\u003e\u003cem\u003eBasic operational evaluation of bending, rotation, and grasping functions\u003c/em\u003e\u003c/strong\u003e\u003c/p\u003e\n\u003cp\u003eEach axis was actuated using motor drive to perform shaft rotation, two-axis bending, and grasping motions, and it was confirmed that all functions operated smoothly (see Additional file 1). Figure 13 shows photographs of the instrument during bending actuation in the pitch and yaw directions, confirming the effectiveness of the bending function. Figure 14 shows the relationship between the drive pulley rotation angle and the resulting bending angle. As the drive pulley rotation angle increases, the bending angle increases approximately linearly; however, hysteresis of approximately \u0026plusmn;5\u0026deg; is observed, which is presumed to be caused by friction between the wires and the disk holes. This level of hysteresis is considered acceptable for preliminary validation; however, further reduction through surface finishing, chamfering or lubrication will be investigated in future work. Figure 15 shows photographs confirming the grasping function. Although the opening angle is relatively small, approximately 10\u0026deg;, it is confirmed that the grasping function operates effectively. In the present evaluation of the microgripper function, quantitative measurement of the grasping force was beyond the scope of this study. Instead, the evaluation focused on deformation in the grasped state and the confirmation of stable grasping behavior. Quantitative evaluation of grasping force will be addressed in future work.\u003c/p\u003e\n\u003cp\u003e\u003cstrong\u003e\u003cem\u003eEvaluation of shaft deflection and whirling during operation\u003c/em\u003e\u003c/strong\u003e\u003c/p\u003e\n\u003cp\u003eAs analyzed in the previous section, shaft deflection due to buckling induced by wire tension is a potential concern associated with miniaturization. The following describes the evaluation results of shaft deflection during each type of operation. Shaft deflection during bending operation was measured from Fig. 14, and shaft deflection during grasping operation was measured from Fig. 15. For shaft whirling during rotation, videos were recorded while rotating the shaft from 0\u0026deg; to \u0026plusmn;160\u0026deg; at a rotational speed of 10\u0026deg;/s under two conditions: a straight configuration with a bending angle of 0\u0026deg;, and a bent configuration with \u0026minus;45\u0026deg; pitch bending (about the x-axis). From these videos, the states exhibiting the maximum whirling displacement of the shaft were extracted. Figure 16 shows the maximum whirling condition observed during shaft rotation. Table 3 summarizes the shaft deflection and whirling displacement during bending, grasping, and shaft rotation operations.\u003c/p\u003e\n\u003cp\u003eWhile the deflection during bending and grasping operations remains below approximately 0.06 mm, it is confirmed that shaft whirling with a maximum displacement amplitude of 0.1\u0026ndash;0.23 mm occurs during shaft rotation. Comparing the straight and bent configurations, the total whirling amplitude is smaller in the bent configuration. Possible causes of the observed whirling include changes in wire tension associated with shaft rotation and bending actuation, as well as fabrication and assembly tolerances of components related to the shaft rotation axis. Identifying the specific causes requires measurements under various conditions, including three-dimensional measurements rather than single-direction observations and evaluation under different initial pretension settings; therefore, this issue is considered a subject for future work. From the perspective of future clinical application, shaft whirling of 0.1\u0026ndash;0.23 mm during shaft rotation may have a limited impact on surgical procedures that do not involve changes in instrument orientation; however, when orientation changes are involved, the effects on operability and task performance must be carefully evaluated. In human-in-the-loop scenarios, such whirling is not expected to constitute a critical issue, as the surgeon can naturally compensate for the resulting deviations. Furthermore, for autonomous or semi-autonomous targeting, real-time compensation may be feasible, provided that the elastic deformation behavior is sufficiently consistent and predictable. Based on the above basic operational evaluation of the bending, rotation, and grasping functions, it was confirmed that the newly designed 3-DOF robotic instrument with a microgripper and an outer diameter of 0.7 mm has no critical issues from a mechanical design and prototyping perspective. Although the present results demonstrate the fundamental feasibility of the proposed ultra-fine robotic instrument, the evaluation was conducted under limited experimental conditions and is based on N = 1 measurements. Further systematic evaluation under diverse conditions and with larger sample sizes is required to fully characterize the performance and robustness of the system.\u003c/p\u003e"},{"header":"Conclusion","content":"\u003cp\u003eThis paper presented the design and miniaturization of an ultra-fine multi-DOF robotic instrument for ophthalmic MIMS. The proposed instrument integrates a 2-DOF bending mechanism, a shaft rotation axis, and a microgripper within an outer diameter of 0.7 mm by introducing a surface-constrained (V-type) disk-stacked bending mechanism. Design considerations for miniaturization were addressed through (i) improved manufacturability and assembly efficiency of the disk elements, (ii) evaluation of wire passability through reduced guide-hole diameters, and (iii) quantitative assessment of shaft stiffness and Euler buckling under wire pretension and a conservative lateral tip load of 10 mN.\u003c/p\u003e \u003cp\u003eFull-scale prototypes of both 0.9-mm-OD and 0.7-mm-OD instruments were fabricated and assembled, and basic bending, rotation, and grasping functions were experimentally verified. The 0.7-mm-OD instrument exhibited smooth pitch/yaw bending and grasping motion with a bending hysteresis of approximately\u0026thinsp;\u0026plusmn;\u0026thinsp;5\u0026deg;. Shaft deflection during bending and grasping remained below approximately 0.06 mm, while whirling during shaft rotation reached displacement amplitudes of 0.1\u0026ndash;0.23 mm, indicating that further refinement of shaft rotation components and tension management will be beneficial for highly precise orientation control.\u003c/p\u003e \u003cp\u003eOverall, the results demonstrate the feasibility of an ultra-fine robotic instrument with distal 3-DOF orientation and an end-effector suitable for ophthalmic MIMS, and clarify key design trade-offs among miniaturization, stiffness, wire routing, pretension, and stability. The insights obtained in this study provide a foundation for future extensions, including more comprehensive experimental validation, optimization of rotational stability, and integration into teleoperated ophthalmic robotic systems.\u003c/p\u003e"},{"header":"Abbreviations","content":"\u003cp\u003eMinimally invasive surgery \u0026nbsp; \u0026nbsp;MIS\u003c/p\u003e\n\u003cp\u003eMinimally invasive microsurgery \u0026nbsp; \u0026nbsp;MIMS\u003c/p\u003e\n\u003cp\u003eDegree-of-freedom \u0026nbsp; \u0026nbsp;DOF\u003c/p\u003e\n\u003cp\u003eRemote center of motion \u0026nbsp; \u0026nbsp;RCM\u003c/p\u003e"},{"header":"Declarations","content":"\u003cp\u003e\u003cstrong\u003eAuthor contributions\u003c/strong\u003e\u003c/p\u003e\n\u003cp\u003eM.J. led the conceptual planning and mechanical design of the robotic instrument, planned and conducted the verification experiments, performed the data analysis, and drafted the initial manuscript. I.I. played a key role in defining the specifications and requirements for the conceptual design. R.N. contributed to the development of the control system, including the design and implementation of the control software. All authors revised the manuscript and reviewed and approved the final version.\u003c/p\u003e\n\u003cp\u003e\u003cstrong\u003eAcknowledgments\u003c/strong\u003e\u003c/p\u003e\n\u003cp\u003eThe authors acknowledge the support of the Japan Keirin Autorace Foundation (JKA).\u003c/p\u003e\n\u003cp\u003e\u003cstrong\u003eCompeting interests\u003c/strong\u003e\u003c/p\u003e\n\u003cp\u003eThe authors declare that they have no competing interests.\u003c/p\u003e\n\u003cp\u003e\u003cstrong\u003eAvailability of data and materials\u003c/strong\u003e\u003c/p\u003e\n\u003cp\u003eNot applicable.\u003c/p\u003e\n\u003cp\u003e\u003cstrong\u003eEthics approval and consent to participate\u003c/strong\u003e\u003c/p\u003e\n\u003cp\u003eNot applicable.\u003c/p\u003e\n\u003cp\u003e\u003cstrong\u003eFunding\u003c/strong\u003e\u003c/p\u003e\n\u003cp\u003eThis research was supported by the subsidy program (individual research) of the Japan Keirin Autorace Foundation (JKA).\u003c/p\u003e"},{"header":"References","content":"\u003col\u003e\n\u003cli\u003eIntuitive Surgical, Inc., \u0026ldquo;da Vinci Surgical System.\u0026rdquo; Available: https://www.intuitive.com. Accessed: Jan. 13, 2026.\u003c/li\u003e\n\u003cli\u003eM. S. Khan, M. R. Elhage, A. Challacombe, et al., \u0026ldquo;Technical review of the da Vinci surgical telemanipulator,\u0026rdquo; Int. J. Med. Robot. Comput. Assist. Surg., vol. 8, no. 4, pp. 467\u0026ndash;476, 2012, doi:10.1002/rcs.1468.\u003c/li\u003e\n\u003cli\u003eP. Probst, J. Keller, L. M\u0026uuml;ller, et al., \u0026ldquo;A review of the role of robotics in surgery: to da Vinci and beyond,\u0026rdquo; Frontiers in Surgery, vol. 10, Art. no. 1198574, 2023, doi:10.3389/fsurg.2023.1198574.\u003c/li\u003e\n\u003cli\u003eMicrosure B.V., \u0026ldquo;MUSA-2, MUSA-3.\u0026rdquo; Available: https://microsure.nl/. Accessed: Jan. 13, 2026.\u003c/li\u003e\n\u003cli\u003eMedical Microinstruments S.p.A., \u0026ldquo;The Symani\u0026reg; Surgical System.\u0026rdquo; Available: https://www.mmimicro.com/. Accessed: Jan. 13, 2026.\u003c/li\u003e\n\u003cli\u003eE. Vander Poorten et al., \u0026ldquo;Robotic retinal surgery,\u0026rdquo; in Handbook of Robotic and Image-Guided Surgery. Elsevier, 2019, ch. 36, pp. 627\u0026ndash;672.\u003c/li\u003e\n\u003cli\u003eY.-Q. Chen et al., \u0026ldquo;Cooperative robot assistant for vitreoretinal microsurgery: Development of the RVRMS and feasibility studies in an animal model,\u0026rdquo; Graefe\u0026rsquo;s Arch. Clin. Exp. Ophthalmol., vol. 255, pp. 1167\u0026ndash;1171, 2017.\u003c/li\u003e\n\u003cli\u003eG. U. M\u0026uuml;ller et al., \u0026ldquo;Robotic retinal surgery impacts on scleral forces: An in vivo study,\u0026rdquo; Transl. Vis. Sci. Technol., vol. 9, no. 10, 2020.\u003c/li\u003e\n\u003cli\u003eF. Ullrich et al., \u0026ldquo;Mobility experiments with microrobots for minimally invasive intraocular surgery,\u0026rdquo; Invest. Ophthalmol. Vis. Sci., vol. 54, pp. 2853\u0026ndash;2863, 2013.\u003c/li\u003e\n\u003cli\u003eA. G. Andy et al., \u0026ldquo;In-human robot-assisted retinal vein cannulation: A world first,\u0026rdquo; Ann. Biomed. Eng., vol. 46, no. 10, pp. 1676\u0026ndash;1685, 2018.\u003c/li\u003e\n\u003cli\u003eT. L. Edwards et al., \u0026ldquo;First-in-human study of the safety and viability of intraocular robotic surgery,\u0026rdquo; Nat. Biomed. Eng., vol. 2, pp. 649\u0026ndash;656, 2018.\u003c/li\u003e\n\u003cli\u003eF.-Y. Lin, C. Bergeles, and G.-Z. Yang, \u0026ldquo;Biometry-based concentric tube robot for vitreoretinal surgery,\u0026rdquo; in Proc. IEEE EMBS, 2015.\u003c/li\u003e\n\u003cli\u003eP. J. Swaney et al., \u0026ldquo;Design, fabrication, and testing of a needle-sized wrist for surgical instruments,\u0026rdquo; J. Med. Devices, vol. 11, no. 1, 014501, 2017.\u003c/li\u003e\n\u003cli\u003eT. Zhang, Z. Ping, and S. Zuo, \u0026ldquo;Miniature continuum manipulator with 3-DOF force sensing for retinal microsurgery,\u0026rdquo; J. Mech. Robot., vol. 13, no. 4, 041002, 2021.\u003c/li\u003e\n\u003cli\u003eN. E. Pacheco et al., \u0026ldquo;Beyond constant curvature: A new mechanics model for unidirectional notched-tube continuum wrists,\u0026rdquo; J. Med. Robot. Res., vol. 6, nos. 1\u0026ndash;2, 2140004, 2021.\u003c/li\u003e\n\u003cli\u003eC. J. Nwafor et al., \u0026ldquo;The Caturo: A submillimeter diameter glass concentric tube robot with high curvature,\u0026rdquo; 2023.\u003c/li\u003e\n\u003cli\u003eT. L. Bruns et al., \u0026ldquo;A modular, multi-arm concentric tube robot system with application to transnasal surgery for orbital tumors,\u0026rdquo; Int. J. Robot. Res., vol. 40, nos. 2\u0026ndash;3, pp. 521\u0026ndash;533, 2021.\u003c/li\u003e\n\u003cli\u003eX. He et al., \u0026ldquo;IRIS: Integrated robotic intraocular snake,\u0026rdquo; in Proc. IEEE ICRA, 2015, pp. 1764\u0026ndash;1769.\u003c/li\u003e\n\u003cli\u003eN. Simaan et al., \u0026ldquo;Design and integration of a telerobotic system for minimally invasive surgery of the throat,\u0026rdquo; Int. J. Robot. Res., vol. 28, no. 9, pp. 1134\u0026ndash;1153, 2009.\u003c/li\u003e\n\u003cli\u003eL. Yan et al., \u0026ldquo;SnakeRaven: Teleoperation of a 3D-printed snake-like manipulator integrated to the RAVEN II surgical robot,\u0026rdquo; in Proc. IEEE/RSJ IROS, 2021.\u003c/li\u003e\n\u003cli\u003eY.-J. Kim et al., \u0026ldquo;A stiffness-adjustable hyperredundant manipulator using a variable neutral-line mechanism for minimally invasive surgery,\u0026rdquo; IEEE Trans. Robot., vol. 30, no. 2, pp. 382\u0026ndash;395, 2014.\u003c/li\u003e\n\u003cli\u003eM. Jinno and I. Iordachita, \u0026ldquo;Improved integrated robotic intraocular snake: Analyses of the kinematics and drive mechanism of the dexterous distal unit,\u0026rdquo; J. Med. Robot. Res., vol. 6, nos. 1\u0026ndash;2, 2140001, 2021.\u003c/li\u003e\n\u003cli\u003eM. Jinno et al., \u0026ldquo;An integrated high-dexterity cooperative robotic assistant for intraocular micromanipulation,\u0026rdquo; in Proc. IEEE ICRA, 2021.\u003c/li\u003e\n\u003cli\u003eM. Jinno and I. Iordachita, \u0026ldquo;Microgripper using flexible wire hinge for robotic intraocular snake,\u0026rdquo; in Proc. IEEE ICRA, 2022.\u003c/li\u003e\n\u003cli\u003eM. Jinno, R. Nonoyama, and I. Iordachita, \u0026ldquo;Multi-degree-of-freedom bending mechanisms and drive mechanisms for robot-assisted minimally invasive microsurgery,\u0026rdquo; in Proc. IEEE/ASME AIM, 2025.\u003c/li\u003e\n\u003cli\u003eJ. Song, C. Gonenc, J. Guo, I. Iordachita, \u0026ldquo;Intraocular Snake Integrated with the Steady-Hand Eye Robot for Assisted Retinal Microsurgery,\u0026rdquo; in Proc. IEEE ICRA, 2017.\u003c/li\u003e\n\u003cli\u003eS. Sunshine et al., \u0026ldquo;A force-sensing microsurgical instrument that detects forces below human tactile sensation,\u0026rdquo; Retina, vol. 33, no. 1, pp. 200\u0026ndash;206, 2013, doi:10.1097/IAE.0b013e3182625d2b.\u003c/li\u003e\n\u003cli\u003eP. K. Gupta, P. S. Jensen, and E. de Juan Jr., \u0026ldquo;Surgical forces and tactile perception during retinal microsurgery,\u0026rdquo; in Proc. MICCAI, LNCS vol. 1679, Springer, 1999. \u003c/li\u003e\n\u003c/ol\u003e"},{"header":"Tables","content":"\u003cp\u003eTables 1 to 3 are available in the Supplementary Files section.\u003c/p\u003e"}],"fulltextSource":"","fullText":"","funders":[],"hasAdminPriorityOnWorkflow":false,"hasManuscriptDocX":true,"hasOptedInToPreprint":true,"hasPassedJournalQc":"","hasAnyPriority":false,"hideJournal":false,"highlight":"","institution":"","isAcceptedByJournal":true,"isAuthorSuppliedPdf":false,"isDeskRejected":"","isHiddenFromSearch":false,"isInQc":false,"isInWorkflow":false,"isPdf":false,"isPdfUpToDate":true,"isWithdrawnOrRetracted":false,"journal":{"display":true,"email":"[email protected]","identity":"robomech-journal","isNatureJournal":false,"hasQc":true,"allowDirectSubmit":false,"externalIdentity":"robo","sideBox":"Learn more about [ROBOMECH Journal](http://robomechjournal.springeropen.com/)","snPcode":"40520","submissionUrl":"https://submission.nature.com/new-submission/40520/3","title":"ROBOMECH Journal","twitterHandle":"@SpringerEng","acdcEnabled":true,"dfaEnabled":true,"editorialSystem":"stoa","reportingPortfolio":"BMC/SO AJ","inReviewEnabled":true,"inReviewRevisionsEnabled":true},"keywords":"Medical robotics, Retinal microsurgery, Ultra-fine robotic instrument, Disk-stacked bending mechanism, Mechanical design","lastPublishedDoi":"10.21203/rs.3.rs-8771621/v1","lastPublishedDoiUrl":"https://doi.org/10.21203/rs.3.rs-8771621/v1","license":{"name":"CC BY 4.0","url":"https://creativecommons.org/licenses/by/4.0/"},"manuscriptAbstract":"\u003cp\u003eMinimally invasive surgery (MIS) has transformed surgical practice by reducing patient trauma and improving postoperative outcomes. In laparoscopic surgery, these benefits have been further enhanced by the clinical adoption of teleoperated robotic systems, most notably the da Vinci Surgical System, which provides improved dexterity, motion scaling, and ergonomics in confined environments. As surgical robotics advances, its application is expected to extend beyond conventional MIS to microsurgical procedures requiring levels of precision and stability beyond those achievable manually. However, the clinical adoption of robotic assistance in microsurgery remains limited, particularly for minimally invasive procedures in highly constrained workspaces. Teleoperated leader\u0026ndash;follower robotic architectures offer a promising solution for robot-assisted minimally invasive microsurgery (MIMS) by enabling precise motion scaling and tremor suppression while preserving intuitive surgeon control.\u003c/p\u003e \u003cp\u003eOphthalmic MIMS requires dexterous manipulation within an extremely confined intraocular workspace under millinewton-level interaction forces. Although snake-like and continuum instruments have been explored to improve access and distal dexterity, achieving multi-degree-of-freedom (DOF) motion within a submillimeter outer diameter remains challenging. These challenges stem from inherent trade-offs among bending range, shaft stiffness, wire routing, pretension, and buckling stability. This work presents the design and miniaturization of an ultra-fine multi-DOF robotic instrument for vitreoretinal surgery.\u003c/p\u003e \u003cp\u003eThe proposed instrument integrates 2-DOF distal bending (pitch and yaw), shaft rotation (roll), and a microgripper within a 0.7 mm outer diameter. To support miniaturization while maintaining manufacturability and structural integrity, a novel surface-constrained, V-type disk-stacked bending mechanism is introduced. Wire passability through reduced-diameter guide holes is geometrically verified at maximum disk tilt, and shaft stiffness and Euler buckling are analyzed using second-moment-of-area models under a conservative 10 mN lateral tip load.\u003c/p\u003e \u003cp\u003ePrototypes with outer diameters of 0.9 mm and 0.7 mm were fabricated and tested. The 0.7 mm instrument demonstrated smooth pitch\u0026ndash;yaw bending and reliable grasping, with bending hysteresis of approximately\u0026thinsp;\u0026plusmn;\u0026thinsp;5\u0026deg;. Shaft deflection during bending and grasping remained below 0.06 mm, while rotational whirling produced displacement amplitudes of 0.1\u0026ndash;0.23 mm. These results highlight key design trade-offs and provide experimentally validated guidelines for the development of ultra-fine robotic instruments for ophthalmic MIMS.\u003c/p\u003e","manuscriptTitle":"Design and Miniaturization of an Ultra-Fine Multi-Degree-of-Freedom Robotic Instrument for Ophthalmic Minimally Invasive Microsurgery","msid":"","msnumber":"","nonDraftVersions":[{"code":1,"date":"2026-02-13 11:50:32","doi":"10.21203/rs.3.rs-8771621/v1","editorialEvents":[{"type":"communityComments","content":0},{"type":"decision","content":"Revision requested","date":"2026-03-31T05:53:28+00:00","index":"","fulltext":""},{"type":"editorInvitedReview","content":"","date":"2026-03-31T05:35:11+00:00","index":"hide","fulltext":""},{"type":"reviewerAgreed","content":"296340347917185956985235893020219169414","date":"2026-03-12T04:31:59+00:00","index":"hide","fulltext":""},{"type":"editorInvitedReview","content":"","date":"2026-03-09T15:47:30+00:00","index":"hide","fulltext":""},{"type":"reviewerAgreed","content":"117511420681225988721473687921260176025","date":"2026-02-09T13:20:43+00:00","index":"hide","fulltext":""},{"type":"reviewersInvited","content":"","date":"2026-02-07T23:16:33+00:00","index":"","fulltext":""},{"type":"editorAssigned","content":"","date":"2026-02-07T23:06:12+00:00","index":"","fulltext":""},{"type":"checksComplete","content":"","date":"2026-02-06T13:08:10+00:00","index":"","fulltext":""},{"type":"submitted","content":"ROBOMECH Journal","date":"2026-02-03T05:46:49+00:00","index":"","fulltext":""}],"status":"published","journal":{"display":true,"email":"[email protected]","identity":"robomech-journal","isNatureJournal":false,"hasQc":true,"allowDirectSubmit":false,"externalIdentity":"robo","sideBox":"Learn more about [ROBOMECH Journal](http://robomechjournal.springeropen.com/)","snPcode":"40520","submissionUrl":"https://submission.nature.com/new-submission/40520/3","title":"ROBOMECH Journal","twitterHandle":"@SpringerEng","acdcEnabled":true,"dfaEnabled":true,"editorialSystem":"stoa","reportingPortfolio":"BMC/SO AJ","inReviewEnabled":true,"inReviewRevisionsEnabled":true}}],"origin":"","ownerIdentity":"7068072a-3208-4dec-b4bb-1fcacbcacd3d","owner":[],"postedDate":"February 13th, 2026","published":true,"recentEditorialEvents":[],"rejectedJournal":[],"revision":"","amendment":"","status":"under-review","subjectAreas":[],"tags":[],"updatedAt":"2026-05-15T12:39:16+00:00","versionOfRecord":[],"versionCreatedAt":"2026-02-13 11:50:32","video":"","vorDoi":"","vorDoiUrl":"","workflowStages":[]},"version":"v1","identity":"rs-8771621","journalConfig":"researchsquare"},"__N_SSP":true},"page":"/article/[identity]/[[...version]]","query":{"redirect":"/article/rs-8771621","identity":"rs-8771621","version":["v1"]},"buildId":"XKTyCvWXoU3ODBz1xrDgd","isFallback":false,"isExperimentalCompile":false,"dynamicIds":[84888],"gssp":true,"scriptLoader":[]}