High Damping Passive Launch Vibration Isolation System Using Superelastic SMA with Multilayered Viscous Lamina
High Damping Passive Launch Vibration Isolation System Using Superelastic SMA with Multilayered...
Park, Yeon-Hyeok;Kwon, Seong-Cheol;Koo, Kyung-Rae;Oh, Hyun-Ung
2021-07-26 00:00:00
aerospace Article High Damping Passive Launch Vibration Isolation System Using Superelastic SMA with Multilayered Viscous Lamina 1 2 2 1 , Yeon-Hyeok Park , Seong-Cheol Kwon , Kyung-Rae Koo and Hyun-Ung Oh * Space Technology Synthesis Laboratory, Department of Smart Vehicle System Engineering, Chosun University, 375 Seosuk-dong, Dong-gu, Gwangju 501-759, Korea; wkfjf6043@chosun.kr Mechatronics Groups, Hanwha Systems, 491-23, Gyeonggidong-ro, Namsa-myeon, Cheoin-gu, Yongin-si 17121, Korea; seongcheol.kwon@hanwha.com (S.-C.K.); kr.koo@hanwha.com (K.-R.K.) * Correspondence: ohu129@chosun.ac.kr Abstract: Whole-spacecraft launch-vibration isolation systems are attractive for achieving the goal of better, faster, cheaper, and lighter small satellites by reducing the design-load and vibration-test specifications for on-board components. In this study, a three-axis passive launch-vibration isolation system, based on superelastic shape memory alloy (SMA) technology, was developed to significantly attenuate the dynamic launch loads transmitted to a small satellite. This provides a superior damping characteristic, achieved by superelastic SMA blades stiffened by multilayered thin plates with viscous lamina adhesive layers of acrylic tape. The basic characteristics of the proposed isolation system with various numbers of viscoelastic multilayers were obtained through a static load test. In addition, the effectiveness of the design was validated through a launch environment simulating sine and random vibration tests. Keywords: superelastic shape memory alloy; whole-spacecraft vibration isolation; small satellite; launch environment Citation: Park, Y.-H.; Kwon, S.-C.; Koo, K.-R.; Oh, H.-U. High Damping Passive Launch Vibration Isolation System Using Superelastic SMA with 1. Introduction Multilayered Viscous Lamina. Aerospace 2021, 8, 201. https:// The start of the New Space paradigm has changed the development philosophy of the doi.org/10.3390/aerospace8080201 worldwide space-engineering field. New Space refers to the recent commercialization of the space sector, which is mainly led by private industries, rather than government-funded Academic Editor: Rosario Pecora organizations. These industries are driving a better, faster, cheaper, and lighter space- development paradigm [1–4]. One of the major accelerators of the New Space paradigm Received: 15 June 2021 is the emergence of small satellite constellations. The large-volume production of small Accepted: 24 July 2021 satellites, based on a standardized satellite platform and commercially available on-board Published: 26 July 2021 hardware, has reduced the development cost and schedule. In addition, the total launch cost can be reduced as multiple small satellites are launched together. These advantages Publisher’s Note: MDPI stays neutral make the small-satellite platform attractive for various challenging missions that require with regard to jurisdictional claims in high temporal system performance, e.g., real-time remote sensing, global internet services, published maps and institutional affil- and high-speed communications [4]. iations. All satellites undergo various dynamic loads during the launch phase. These loads in- volve steady-state acceleration caused by the engine thrust, sinusoidal vibration caused by the engine cutoff, and a self-excited vibration, called the pogo effect, from the combustion instability of the launcher. Random vibrations are caused by the noise of the thrust, and me- Copyright: © 2021 by the authors. chanical shock is induced by the separation of the launcher stages and spacecraft. Because Licensee MDPI, Basel, Switzerland. these dynamic loads are extremely severe and complex, they are one of the major factors This article is an open access article that cause satellites and components to malfunction [5]. Therefore, a proper structural distributed under the terms and design is essential to ensure the structural safety of the satellite in the launch environment. conditions of the Creative Commons Two approaches are used to enhance the structural safety of satellites during the Attribution (CC BY) license (https:// design phase. One traditional approach is to design the spacecraft structure such that it creativecommons.org/licenses/by/ has sufficient strength and stiffness to endure mechanical loading. However, this approach 4.0/). Aerospace 2021, 8, 201. https://doi.org/10.3390/aerospace8080201 https://www.mdpi.com/journal/aerospace Aerospace 2021, 8, 201 2 of 14 faces the technical issue on limitation in minimizing the mass and volume of the satellite, especially in small satellites with design limitations of their volume and mass. These factors are related to the available number of satellites that can be launched within the lift capability of the launcher and, therefore, directly lead to an increase in launch costs. The other method involves reducing the launch loads transmitted to the satellite by applying a whole-spacecraft vibration isolator (WSVI). The WSVI is achieved by implementing a low-stiffness and high-damping capability, compared to the conventional concept, which is rigidly mounted between the satellite and launch adaptor. This makes it possible to effectively reduce the mass and volume of the satellite by minimizing the design load of vibration-sensitive components. In addition, it can contribute to reducing the satellite’s development cost and schedule by optimizing the conventional verification process, which accounts for a large portion of the time and cost of satellite-development programs. In addition, the on-ground test process can be optimized, e.g., simplifying and skipping the test phases of the subsystem-level verification. For example, verification can be performed at a higher system level of the integrated test because using the WSVI greatly reduces the potential technical risks to the vibration-sensitive components at the subsystem level. Another possible test schedule mitigation might also be expected on the notching process in the launch-vibration tests because creating a proper notched input level takes a large portion of the time. These potential advantages of WSVIs are attractive for the New Space based development trend to achieve the goal of cost-effective small satellite development. Several previous studies have been conducted to realize whole-spacecraft vibration isolators. Johnson et al. [6] proposed WSVI systems called SoftRide UniFlex and SoftRide MultiFlex to realize launch-vibration isolation for the MINOTAUR/JAWSAT program. SoftRide UniFlex consisted of a titanium flexure and a viscoelastic material to reduce the dynamic loads acting on the satellite along the launcher axial direction. To isolate both axial and lateral excitations from the launcher, SoftRide MultiFlex consisted of a pair of UniFlex isolators integrated with each other through a central post. The vibration in the launcher axial direction was isolated in the same manner as the UniFlex. Isolation in the lateral directions was achieved by vibrational-energy dissipation from the shear deformation of the constrained layers, as the dynamic bending of the flexure occurred. To enhance the performance of SoftRide MultiFlex, SoftRide OmniFlex was developed by Johal et al. [7]. The major difference from the previous version was that the constrained layer was installed at the sides of the titanium flexure. This enabled the isolator to be more compact in volume, as well as enhancing the damping capability. Jun et al. [8] proposed an isolation system using two serially connected flexure elements and a shear damping unit laminated with a metal plate and viscoelastic material. Isolation in the axial and lateral directions was achieved by the bending behavior of the titanium flexure and the shearing behavior of the damping unit, respectively. Mastroddi et al. [9] proposed a multi-frequency dynamic absorber using a two-spring mechanism arranged in each axial and lateral direction. These mechanisms are connected with the oscillating mass for tuning the specific target frequency for the spacecraft during the launch phase. Rittweger et al. [10] proposed an active payload adaptor compliant with the Ariane 5 space-launch vehicle, which was able to reduce the interface loads to the payload in the 5–100 Hz low frequency domain by more than a factor of four. In actualizing a novel high damping three-axis passive launch vibration isolation for small satellite, we focused on the following two technical aspects. One is using the superelasticity of the shape memory alloy (SMA) material, and the other is applying multilayered thin plates with viscous lamina tapes to the SMA blades. The superelasticity is a unique characteristic of SMA material occurred by stress-induced phase transformation from the austenite to the martensite phase when the material is at above austenite finish temperature. It can be deformed considerably without being plastically deformed and recovers its original shape upon unloading. This makes it possible to ensure the structural safety of a satellite under the launch environment, even if the satellite is supported by Aerospace 2021, 8, 201 3 of 14 a low-stiffness SMA application. In addition, this characteristic is associated with large hysteretic damping, owing to the phase transformation [11]. The effectiveness of using superelastic SMA for vibration isolation in space applications has been demonstrated in several previous studies [12,13]. For example, Kwon et al. [12] proposed a superelastic SMA gear wheel that applied a two-axis gimbal-type X-band antenna to enhance the micro-jitter isolation capability without undergoing plastic defor- mation under excessive loading conditions. Kwon et al. [13] also proposed a blade-type cooler vibration isolator using a superelastic SMA material as a technical solution over ordinary titanium material. Then, the effectiveness of the superelastic SMA blade design was experimentally assessed and compared with that of an isolator made of titanium. If the superelastic SMA is applied to the WSVI, it is possible to ensure the structural safety of the satellite in a launch environment, while implementing a low-stiffness isolator. However, the SMA exhibits effective hysteretic damping only when the stress exceeds the critical point at which a phase transformation occurs. This means that an isolator with SMA material may exhibit insufficient damping performance, if the deformation is not sufficiently large. Therefore, in this study, we proposed to apply multilayered thin plates with viscous lamina tapes to maximize the damping capability, even under a relatively small deformation. The effectiveness of viscoelastic multilayered thin plates in enhancing the damping capability has been investigated in many studies [14–17]. Minesugi et al. [14] proposed damping mechanisms of polyimide tape with a viscous lamina to reduce the vibration transmitted to the battery panel of the MUSES-A satellite. The experimental results showed that a five times higher damping performance was obtained when viscous lamina was applied, as compared to that without the lamina. Bhattarai et al. [15] proposed a highly damped deployable solar-panel module using a multilayered stiffener with viscoelastic acrylic tape. The design effectiveness of the solar-panel module was validated through a launch-vibration test. Park et al. [16] proposed a high damping printed circuit board (PCB) with multi- layered viscoelastic acrylic tapes for use in wedge lock applications. This PCB concept was effective in increasing the fatigue life of electronic packages, owing to the highly increased damping capability, as well as minimizing the volume and mass of the elec- tronics. Stoudt et al. [17] hypothesized that using nanoscale multilayer coatings could significantly increase the fatigue durability. To confirm this hypothesis, cyclic-loading fatigue tests were performed with Cu-coated films and different surface treatments, in- cluding a nanoscale Cu-Ni multilayer. The results of the fatigue tests showed that the Cu-Ni multilayer film had a fatigue life more than six times greater than that of the other general films under cyclic-loading conditions because the slip between each layer acted as a stress-energy–dissipation mechanism. In this study, the basic characteristics of the passive WSVI with various number of interlaminated layers on the superplastic SMA blade, designed for vibration isolation of 40 kg class satellite, were obtained through static load tests. In addition, to validate the effectiveness of the design in terms of the launch-load attenuation, sine and random vibration tests were performed using a mass-simulating dummy satellite. These test results demonstrated that the proposed WSVI is effective for achieving a novel design goal of both superelastic and high damping capability. 2. WSVI Design Description Figure 1a,b show isometric and internal views of the proposed WSVI. This WSVI was developed to reduce the launch loads above the range of the 28 Hz target cut-off frequency for a 40 kg class satellite. The design strategy of the isolator involves implementing a high damping capability by applying a superelastic SMA blade with a multilayered thin plate with viscous lamina tapes. The SMA material can be applied because it provides superelastic characteristics, high damping capability, and fatigue durability, compared to ordinary metal materials [12]. Superelastic behavior is known to have complete re- Aerospace 2021, 8, x FOR PEER REVIEW 4 of 14 with viscous lamina tapes. The SMA material can be applied because it provides supere- Aerospace 2021, 8, 201 4 of 14 lastic characteristics, high damping capability, and fatigue durability, compared to ordi- nary metal materials [12]. Superelastic behavior is known to have complete reversibility for strains of up to 10–12% without being plastically deformed, which is a very uncommon feature in ordinary metal materials [11]. Therefore, the SMA blade makes it possible to versibility for strains of up to 10–12% without being plastically deformed, which is a very ensure the structural safety of the low stiffness blade of the isolator under the launch en- uncommon feature in ordinary metal materials [11]. Therefore, the SMA blade makes it vironment. possible to ensure the structural safety of the low stiffness blade of the isolator under the launch environment. (a) (b) Figure 1. Configuration of the Proposed WSVI ((a) Isometric View, (b) Inside View). Figure 1. Configuration of the Proposed WSVI ((a) Isometric View, (b) Inside View). In addition, multilayered thin metal plates with viscous lamina tapes were utilized to In addition, multilayered thin metal plates with viscous lamina tapes were utilized increase the damping capability, which also helped distribute the stress release acting on to increase the damping capability, which also helped distribute the stress release acting the thin SMA blade. The proposed WSVI is mainly composed of SMA blade modules to on the thin SMA blade. The proposed WSVI is mainly composed of SMA blade modules achieve the required target cut-off frequency in the axial and lateral directions, a moving plate, upper and bottom plates, inner and outer brackets, and displacement limiters to limit the movement of the blades to within their allowable deflection range. The stiffness of the vibration isolator for each axis is significantly governed by the SMA blade modules. Aerospace 2021, 8, 201 5 of 14 To achieve the required isolator stiffness, the SMA blade modules for the axial direction are connected between the inner and outer brackets at 120 intervals. The SMA blade modules for the lateral direction are arranged between the moving plate and inner bracket, in the same manner as those of the axial direction. The blade is made of superplastic SMA, and it provides a mechanical interface to attach viscoelastic adhesive tapes with thin constrained layers made of FR-4 material. Two constrained layers were applied to each side of the SMA blade. The physical contact between the constrained layers and adhesive tapes reduces the resultant stress acting on the SMA blade because the viscoelasticity resists additional deformations of the SMA blade. Furthermore, each boundary layer between the adhesive tapes and constrained layers experiences shear deformation when the multilayered SMA blade is deformed. This design contributes to the excellent damping performance and enhanced structural safety of the SMA blades. To limit the deformation of the blades within the allowable range of 5 mm, displacement limiters made of high damping plastic material of Delrin [18], with space heritages are included on the design, which is helpful to dissipate and mitigate launch vibration loads when slip and contact occur between the plastic and plastic materials. The mechanical properties of the SMA materials are summarized in Table 1 [12]. The viscoelastic tape used in this study was 3M966 double-sided acrylic tape (3M) [19], as listed in Table 2. Table 1. Material Properties of the Superelastic SMA. Characteristic Value Martensite Finish Temperature (M , C) 21 Martensite Start Temperature (M , C) 12 Austenite Start Temperature (A , C) 5 Austenite Finish Temperature (A , C) 15 Martensite 75 Young’s Modulus (GPa) Austenite 80 Tensile Strength (MPa) 1300 Elongation at Break (%) 45 Density (g/cm ) 6.45 Poisson’s Ratio () 0.33 Table 2. Specifications of Viscoelastic Adhesive Tape (3M966). Item Specification Type Double-sided Acrylic Tape Thickness (mm) 0.06 58 (20-min Dwell) Adhesion Strength to Steel (N/100 mm) 85 (72-h Dwell) 159 (Ultimate Bond) Outgassing (%, TML/CVCM) 0.93/0.01 A static load test was performed at room temperature to evaluate the basic charac- teristics of the proposed WSVI. Two WSVI cases with different SMA blade thicknesses (1) 0.8 mm and (2) 1.5 mm were used in this test. Figure 2 shows an example of a static load test setup for the lateral axis. The WSVI was connected to the load cell using a grip and a connector, which was mounted on the upper side of the static testing machine. In addition, the Delrin parts of the WSVI were intentionally removed to check the internal status of the blades during the test. Repeated translational loadings of three cycles in the positive and negative directions were applied to the WSVI in each axis to measure its Aerospace 2021, 8, x FOR PEER REVIEW 6 of 14 test setup for the lateral axis. The WSVI was connected to the load cell using a grip and a connector, which was mounted on the upper side of the static testing machine. In addition, Aerospace 2021, 8, 201 6 of 14 the Delrin parts of the WSVI were intentionally removed to check the internal status of the blades during the test. Repeated translational loadings of three cycles in the positive and negative directions were applied to the WSVI in each axis to measure its basic char- basic characteristics. The translational displacements acting on the SMA blade modules in acteristics. The translational displacements acting on the SMA blade modules in Cases (1) Cases (1) and (2) are 4 mm and 3 mm, respectively, with a velocity of 2 mm/min. The and (2) are ±4 mm and ±3 mm, respectively, with a velocity of 2 mm/min. The correspond- corresponding displacement ranges are the estimated maximum displacements derived ing displacement ranges are the estimated maximum displacements derived from the from the launch vibration analysis. launch vibration analysis. Figure 2. Example of the Static Load Test Setup for the Lateral Axis. Figure 2. Example of the Static Load Test Setup for the Lateral Axis. Figur Figure e 3 3 a,b a,b show the l show the load oaddisplacement displacement rel relations ations obta obtained ined f frr om om the the static staticload loadtest test of of the WSVI in the lateral and axial directions for Cases (1) and (2). These test results show the WSVI in the lateral and axial directions for Cases (1) and (2). These test results show that thatthe the e equivalent quivalent stif st fi ness ffness o of Case f Ca(2), se (with 2), w aitblade h a blthickness ade thickn ofess 1.5 o mm, f 1.5is m 1.7 m,times is 1.7 t higher imes than higher tha that ofnCase that of (1), Case with ( a10.8 ), wi mm th a blade. 0.8 mm bl In thead test, e. In the test, the bounda the boundary layers ofry la the thin yers of plates the with viscous lamina tapes did not delaminate, and no plastic deformation was observed on thin plates with viscous lamina tapes did not delaminate, and no plastic deformation was the SMA blade within the tested range of translational loading. Because all subsequent observed on the SMA blade within the tested range of translational loading. Because all curves completely coincided with the initial hysteresis curve, even though the structural subsequent curves completely coincided with the initial hysteresis curve, even though the analysis results from the blades made of titanium and aluminum showed a negative margin structural analysis results from the blades made of titanium and aluminum showed a neg- of safety. Furthermore, the results showed a much larger hysteresis area, which cannot be ative margin of safety. Furthermore, the results showed a much larger hysteresis area, achieved by general metal materials, including superplastic SMA [12]. This is because the which cannot be achieved by general metal materials, including superplastic SMA [12]. slip and friction induced by strong molecular attraction forces between multi laminated This is because the slip and friction induced by strong molecular attraction forces between plates with the viscous lamina tape helped enhance the damping capability of the WSVI, because of the larger shear deformation and strain on the multi- laminated SMA blades. In addition, the WSVI in the axial direction showed a much higher damping characteristic than that in the lateral direction because a 2.5 times larger area of the hysteresis curve was obtained. This can be explained by the fact that the shear deformation of the SMA blades arranged in the axial direction is larger than that in the lateral direction. A structural Aerospace 2021, 8, x FOR PEER REVIEW 7 of 14 multi laminated plates with the viscous lamina tape helped enhance the damping capa- bility of the WSVI, because of the larger shear deformation and strain on the multi- lami- nated SMA blades. In addition, the WSVI in the axial direction showed a much higher damping characteristic than that in the lateral direction because a 2.5 times larger area of the hysteresis curve was obtained. This can be explained by the fact that the shear defor- Aerospace 2021, 8, 201 7 of 14 mation of the SMA blades arranged in the axial direction is larger than that in the lateral direction. A structural analysis of the SMA blade for the axial direction also showed a 2.2 times greater maximum strain than that of the blade for the lateral direction. The test re- sults also showed that the hysteresis area obtained from Case (2) was almost the same or analysis of the SMA blade for the axial direction also showed a 2.2 times greater maximum slightly less than that of Case (1). Moreover, a nonlinear characteristic induced by the strain than that of the blade for the lateral direction. The test results also showed that phase transformation of the superplastic SMA blade was not observed within the tested the hysteresis area obtained from Case (2) was almost the same or slightly less than that displacement range. On the contrary, the stiffness of the WSVI in the axial direction in- of Case (1). Moreover, a nonlinear characteristic induced by the phase transformation of creased the superplastic as the displaceme SMA blade nt incre was not ased observed . This phenom withinenon seems to be mainly r the tested displacement range. elatedOn to the WSVI design because the stiffness in the longitudinal direction of the SMA blade be- the contrary, the stiffness of the WSVI in the axial direction increased as the displacement comes domin increased. This ant phenomenon in determining seems the st toif be fness o mainly f thre elated WSVI to as the the d WSVI isplacement design because increasthe es. Therefore stiffness in , these re the longitudinal sults from the static direction of load the SMA test of blade the WSV becomes I in dominant dicated tha in t determining the design the stiffness of the WSVI as the displacement increases. Therefore, these results from the strategy of applying multi-layered viscous lamina to the SMA blades is a much more dom- static load test of the WSVI indicated that the design strategy of applying multi-layered inant factor in enhancing the damping of the isolator than the inherent damping perfor- mance of viscous lamina superelastic to theSM SMA A. In ad blades ditio is n a, the much multilayered viscous la more dominant factor miin na contributes to enhancing the damping of the isolator than the inherent damping performance of superelastic SMA. In reducing the stress acting on the thin SMA blade, owing to the viscoelasticity of the adhe- addition, the multilayered viscous lamina contributes to reducing the stress acting on the sive tapes [17]. Furthermore, a previous study [12] reported that a superelastic SMA de- thin SMA blade, owing to the viscoelasticity of the adhesive tapes [17]. Furthermore, a veloped for high damping SMA gear showed higher fatigue durability than titanium un- previous study [12] reported that a superelastic SMA developed for high damping SMA der cyclic loadings. Therefore, SMA blades may be useful for launch-vibration isolator gear showed higher fatigue durability than titanium under cyclic loadings. Therefore, SMA applications. blades may be useful for launch-vibration isolator applications. Case 1 Case 2 -400 -800 -1200 -1600 -5 -2.5 0 2.5 5 Displacement (mm) (a) Figure 3. Cont. Load (N) Aerospace 2021, 8, 201 8 of 14 Aerospace 2021, 8, x FOR PEER REVIEW 8 of 14 Case 1 Case 2 -400 -800 -1200 -1600 -5 -2.5 0 2.5 5 Displacement (mm) (b) Figure 3. Static Load Test Results ((a) Lateral Direction, (b) Axial Direction). Figure 3. Static Load Test Results ((a) Lateral Direction, (b) Axial Direction). The estimated values of the equivalent damping z for each case are summarized in eq The estimated values of the equivalent damping 𝜁 for each case are summarized Table 3. The values of z in each direction were obtained by the following equivalent lin- eq in Table 3. The values of 𝜁 in each direction were obtained by the following equivalent earization method, in which the nonlinear stiffness and damping coefficients are translated linearization method, in which the nonlinear stiffness and damping coefficients are trans- into linear ones [20]. lated into linear ones [20]. DE(a ) z (a ) = (1) eq 0 ∆𝐸(𝛼 ) 2pa k eq 𝜁 (𝛼 ) = (1) 2𝜋𝑎 𝑘 where, k is estimated from the linear-curve fitting of the overall slope of the load- eq displacement curve. DE is the area of the closed loop of the hysteresis curve and a where, 𝑘 is estimated from the linear-curve fitting of the overall slope of the load-dis- is the amplitude of the displacement. placement curve. ∆𝐸 is the area of the closed loop of the hysteresis curve and 𝑎 is the amplitude of the displacement. Table 3. Static Load Test Results of the Proposed WSVI. Table 3. Static Load Test Results of the Proposed WSVI. Equivalent Damping Ratio (z ) Equivalent Stiffness (k , N/mm) eq eq Case Equivalent Damping Ratio (𝛇 ) Equivalent Stiffness (𝒌 , N/mm) Axial Lateral Axial Lateral (z-Axis) (x- and y-Axes) (z-Axis) (x- and y-Axes) Case Axial Lateral Axial Lateral (z-Axis) (x- and y-Axes) (z-Axis) (x- and y-Axes) 1 0.095 0.044 219.5 175.8 1 0.095 0.044 219.5 175.8 2 0.087 0.032 386.6 325.4 2 0.087 0.032 386.6 325.4 3. Design Validation Test 3. Design Validation Test To verify the effectiveness of the proposed WSVI design under a launch vibration To verify the effectiveness of the proposed WSVI design under a launch vibration environment, i.e., performance of launch load reduction and structure safety of the WSVI, environment, i.e., performance of launch load reduction and structure safety of the WSVI, sine and random vibration tests were performed at the qualification level. Low Level sine and random vibration tests were performed at the qualification level. Low Level Sine Load (N) 𝒆𝒒 𝒆𝒒 Aerospace 2021, 8, 201 9 of 14 Aerospace 2021, 8, x FOR PEER REVIEW 9 of 14 Sine Sweep (LLSS) tests were performed before and after the vibration test to confirm the Sweep (LLSS) tests were performed before and after the vibration test to confirm the char- characteristic variations of the WSVI. acteristic variations of the WSVI. Figure 4 shows an example of the launch vibration test setup for the WSVI condition Figure 4 shows an example of the launch vibration test setup for the WSVI condition on the z-axis. In the test setup, to achieve a 28 Hz cut-off frequency for a 40 kg class on the z-axis. In the test setup, to achieve a 28 Hz cut-off frequency for a 40 kg class satel- satellite, four WSVIs with the Case (1) design were integrated with a mass simulating lite, four WSVIs with the Case (1) design were integrated with a mass simulating dummy dummy satellite. The dummy satellite was configured as a flat plate-type to implement satellite. The dummy satellite was configured as a flat plate-type to implement the design the design concept of a small synthetic aperture radar (SAR) technology experimental concept of a small synthetic aperture radar (SAR) technology experimental project (S- project (S-STEP) satellite [21]. The S-STEP satellite is an 80 kg class small SAR-satellite that STEP) pr s ovides atellite [ a high 21]. The resolution S-STEP 1sat m e str llit ipmap e is an image. 80 kg cl Itass was sm designed all SAR-sat with ellit ae t flat haplate-type t provides a high resolution 1 m stripmap image. It was designed with a flat plate-type structure for structure for mechanical design simplicity and high dimensional stability in orbit. In the mechanical design simplicity and high dimensional stability in orbit. In the test, the rigid test, the rigid mounted condition was also exposed to the vibration test for comparison with mounted condi the WSVI.tiThe on wa vibration s also exposed to input from the vi thebra vibration tion test f shaker or comp was arison wi obtained th tfr he WSV om the I. r The vibration eference acceler input from th ometer. The e vibration vibration r sh esponses aker wafor s obtained from the referenc each axis of the WSVI wer e acc e obtained elerom- from a three-axis accelerometer placed on the dummy satellite, as shown in Figure 4. The eter. The vibration responses for each axis of the WSVI were obtained from a three-axis qualification levels of the sine and random vibration test specifications applied to the accelerometer placed on the dummy satellite, as shown in Figure 4. The qualification lev- design verification of the WSVI are listed in Tables 4 and 5, and the axes of the test are els of the sine and random vibration test specifications applied to the design verification shown in Figure 4. of the WSVI are listed in Tables 4 and 5, and the axes of the test are shown in Figure 4. Figure Figure 4. 4. Example Example of of the Laun the Launch ch Vibration T Vibration Test est S Setup etup for th for thee zz- -axis. axis. Table 4. Specification of the Sine-vibration Test. Table 4. Specification of the Sine-vibration Test. Item Specification Item Specification Direction Direction x x,, y y,,z z Duration Duration 2 2 oct oct//min min Frequency [Hz] Acceleration [g] Frequency [Hz] Acceleration [g] 5 1 5 1 Profile Profile 15 1.25 15 1.25 100 1.25 100 1.25 Table 5. Specification of the Random-vibration Test. Item Specification Direction x, y, z RMS Acceleration 11.64 grms Aerospace 2021, 8, 201 10 of 14 Table 5. Specification of the Random-vibration Test. Aerospace 2021, 8, x FOR PEER REVIEW 10 of 14 Item Specification Direction x, y, z RMS Acceleration 11.64 g rms Duration 2 min Duration 2 min Frequency [Hz] PSD [G /Hz] Frequency [Hz] PSD [G /Hz] 20 0.014 20 0.014 80 0.044 80 0.044 160 0.07 160 0.07 Profile 640 0.07 Profile 640 0.07 800 0.12 800 0.12 1150 0.12 1150 0.12 1300 0.04 1300 0.04 2000 0.04 2000 0.04 A modal analysis was performed to investigate the dynamic behavior of the dummy A modal analysis was performed to investigate the dynamic behavior of the dummy satellite combined with the WSVI. Figure 5 shows the results of the modal analysis. The satellite combined with the WSVI. Figure 5 shows the results of the modal analysis. The first and second modes at 28 and 30 Hz, respectively, mainly represent the bending mode first and second modes at 28 and 30 Hz, respectively, mainly represent the bending mode of the blade modules for the lateral direction in the x- and y-axes. The third mode at 34 Hz of the blade modules for the lateral direction in the x- and y-axes. The third mode at 34 Hz indicates the bending mode of the blade modules of the WSVI in the axial direction along indicates the bending mode of the blade modules of the WSVI in the axial direction along the z-axis. The fourth mode at 40 Hz indicates the local bending mode of the center of the the z-axis. The fourth mode at 40 Hz indicates the local bending mode of the center of the dummy satellite. These results were used to investigate the launch-vibration test results. dummy satellite. These results were used to investigate the launch-vibration test results. (a) (b) (c) (d) Figure 5. Modal Analysis Results ((a) 1st Mode, (b) 2nd Mode, (c) 3rd Mode, and (d) 4th Mode). Figure 5. Modal Analysis Results ((a) 1st Mode, (b) 2nd Mode, (c) 3rd Mode, and (d) 4th Mode). Figure 6 shows the sine vibration test results for the WSVI during the x-, y-, and Figure 6 shows the sine vibration test results for the WSVI during the x-, y-, and z- z-axis excitation. In the case of the x-axis vibration response, the highest acceleration of the axis excitation. In the case of the x-axis vibration response, the highest acceleration of the dummy satellite with the WSVI was 4.7 g at 28 Hz, which corresponds to the bending mode dummy satellite with the WSVI was 4.7 g at 28 Hz, which corresponds to the bending of the SMA blade modules for the lateral direction of the WSVI along the x-axis. This value mode of the SMA blade modules for the lateral direction of the WSVI along the x-axis. is almost similar to the estimated first eigenfrequency of 28 Hz from the modal analysis This value is almost similar to the estimated first eigenfrequency of 28 Hz from the modal results analys shown is results shown in Figure 5 in F . Inigur addition, e 5. In ad the diti second on, the se response cond respon was followed se was fo by llowe appr doximately by ap- proximately 43 Hz, which was induced by the rotational mode of the SMA blade modules for the lateral direction of the WSVI along the z-axis. The y-axis response shows that the Aerospace 2021, 8, 201 11 of 14 Aerospace 2021, 8, x FOR PEER REVIEW 11 of 14 43 Hz, which was induced by the rotational mode of the SMA blade modules for the lateral direction of the WSVI along the z-axis. The y-axis response shows that the highest highest acceleration of 5.3 g was at 28 Hz. The overall tendency shows characteristics sim- acceleration of 5.3 g was at 28 Hz. The overall tendency shows characteristics similar to the ilar to the x-axis result, owing to the symmetric configuration of the test setup. The z-axis x-axis result, owing to the symmetric configuration of the test setup. The z-axis response response shows the highest acceleration of 11.1 g at 35 Hz, which is similar to the esti- shows the highest acceleration of 11.1 g at 35 Hz, which is similar to the estimated third mated third eigenfrequency of 34 Hz. This response is higher than that of the x- and y- eigenfrequency of 34 Hz. This response is higher than that of the x- and y-axes because axes because of the coupling with the fourth mode of the structural elastic mode of 40 Hz of the coupling with the fourth mode of the structural elastic mode of 40 Hz for the test for the test dummy structure. In addition, the second and third peak responses were fol- dummy structure. In addition, the second and third peak responses were followed at 72 lowed at 72 and 98 Hz from the structural mode of the dummy satellite. From the sine- and 98 Hz from the structural mode of the dummy satellite. From the sine-vibration test vibration test results, it can be seen that the highest acceleration response obtained from results, it can be seen that the highest acceleration response obtained from each of the x-, each of the x-, y-, and z-axes did not exceed the design load of 23 g, derived from the mass y-, and z-axes did not exceed the design load of 23 g, derived from the mass acceleration acceleration curve (MAC) [22]. This indicates that the WSVI is effective for attenuating curve (MAC) [22]. This indicates that the WSVI is effective for attenuating sine vibration sine vibration loads, as the design intends. loads, as the design intends. Input Profile (max. 2.5 g) x-axis Res. (max. 4.7 g) y-axis Res. (max. 5.3 g) z-axis Res. (max. 11.1 g) 0.1 10 100 Frequency (Hz) Figure 6. Sine-Vibration Test Results for x-, y-, and z-axes. Figure 6. Sine-Vibration Test Results for x-, y-, and z-axes. Figure 7a,b show the representative random-vibration test results for the WSVI during Figure 7a,b show the representative random-vibration test results for the WSVI dur- the y- and z-axis excitation. To compare the vibration-reduction capabilities of the WSVI, ing the y- and z-axis excitation. To compare the vibration-reduction capabilities of the the test results obtained from the rigid-mounted condition are also plotted in the Figure. WSVI, the test results obtained from the rigid-mounted condition are also plotted in the The x-axis test result is not shown here, owing to the symmetrical configuration of the WSVI Figure. The x-axis test result is not shown here, owing to the symmetrical configuration of in the x and y planes, as mentioned in the sine-vibration test results. In the case of Figure 7a, the WSVI in the x and y planes, as mentioned in the sine-vibration test results. In the case the first eigenfrequency of the dummy satellite without the WSVI was observed at 975 Hz of Figure 7a, the first eigenfrequency of the dummy satellite without the WSVI was ob- and a maximum acceleration of 23.8 g was observed, with respect to the random test rms served at 975 Hz and a maximum acceleration of 23.8 grms was observed, with respect to input of 11.64 g . However, the maximum acceleration was significantly decreased to rms the random test input of 11.64 grms. However, the maximum acceleration was significantly 2.0 g by applying the WSVI, and the random-vibration response was reduced by a rms decreased to 2.0 grms by applying the WSVI, and the random-vibration response was re- factor of 11.9, comparing with the rigidly mounted condition. The maximum acceleration duced by a factor of 11.9, comparing with the rigidly mounted condition. The maximum response along the z-axis without the WSVI was 114.8 g at 260 Hz. This response is rms acceleration response along the z-axis without the WSVI was 114.8 grms at 260 Hz. This higher than that of the y-axis because the structural elastic modes of the dummy satellite response is higher than that of the y-axis because the structural elastic modes of the are dominant along the z-axis. However, the maximum acceleration of the dummy satellite dummy satellite are dominant along the z-axis. However, the maximum acceleration of the dummy satellite with the WSVI was also significantly reduced to 12.2, as shown in Figure 7b. This indicates that the output response is reduced by a factor of 9.4. Acceleration (g) Aerospace 2021, 8, 201 12 of 14 Aerospace 2021, 8, x FOR PEER REVIEW 12 of 14 with the WSVI was also significantly reduced to 12.2, as shown in Figure 7b. This indicates that the output response is reduced by a factor of 9.4. Input Profile (max. 11.64 g ) rms with WSVI (max. 2.0 g ) rms w/o WSVI (max. 23.8 g ) rms 0.1 0.01 0.001 0.0001 -5 -6 100 1000 Frequency (Hz) (a) 0.1 0.01 0.001 Input Profile (max. 11.64 g ) rms 0.0001 with WSVI (max. 12.2 g ) rms w/o WSVI (max. 114.5 g ) rms -5 100 1000 Frequency (Hz) (b) Figure 7. Random-Vibration Test Results ((a) y-axis, (b) z-axis). Figure 7. Random-Vibration Test Results ((a) y-axis, (b) z-axis). Table Table6 summarizes the 6 summarizes thefirst eigen first eigenfr freequencies quencies of of the dummy sa the dummy satellite tellite wi with th the theWSV WSVI I in in ea each ch aaxis, xis, obta obtained ined through the through theLLSS test LLSS tests s performed before and after ea performed before and after each ch vibra vibration tion test. LLSS tests were performed to validate the structural safety of the SMA multilayered test. LLSS tests were performed to validate the structural safety of the SMA multilayered blade modules by investigating the dynamic responses of the dummy satellite with the WSVI. The results show that the maximum first eigenfrequency shift was within 3.9% PSD Acceleration (g /Hz) PSD Acceleration (g /Hz) Aerospace 2021, 8, 201 13 of 14 blade modules by investigating the dynamic responses of the dummy satellite with the WSVI. The results show that the maximum first eigenfrequency shift was within 3.9% throughout the test event, which was within the 5% criterion of the vibration test [23]. This indicated that no mechanical failures of the WSVI were observed, e.g., the plastic deformation or delamination of the SMA multi-layered blades with viscous lamina. This is because the interlaminated surfaces with the double-sided adhesive are effective in resisting the shear force acting on the multilayered blade with energy-absorption effects. These launch-vibration test results indicate that the SMA multi-layered blade is effective for ensuring the structural safety of the WSVI itself and reducing the transmitted launch loads to the dummy satellite, as the design strategy intended. Table 6. Results of the WSVI’s Low Level Sine Sweep (LLSS) Tests Conducted before and after the Launch-vibration Tests. Excitation Corresponding Axis Frequency Shift Test Status Axis 1st Eigenfrequency (Hz) Difference (%) x After 28.8 - Before 28.2 0.35 Sine Vibration After 28.3 Before 34.2 0.87 After 34.5 Before 28.2 x 2.08 After 28.8 Before 27.9 Random 3.9 Vibration After 26.8 Before 34.4 z 0.86 After 34.7 4. Conclusions In this study, a three-axis passive WSVI was developed to significantly attenuate the dynamic launch loads transmitted to a small satellite. To achieve a high damping capability, the proposed WSVI applied two technical design concepts which are to use the superelasticity of the SMA material and the other is to apply multilayered thin plates with viscous lamina tapes on the SMA blades. The basic characteristics of the proposed WSVI were investigated using static load tests with SMA blades of various thicknesses. The vibra- tion test results to validate the effectiveness of the design showed great launch-vibration isolation performance. From the sine-vibration test results, the highest acceleration re- sponse obtained from each of the x-, y-, and z-axes are reduced 76%, 76%, 52% compared with design load of 23 g. The results of random vibration test showed that the maximum acceleration from each of the x-, y-, and z-axes are reduced 92%, 92%, 90% compared with that of rigid mounted condition. Moreover, the structural safety of the WSVI was within the qualification level of the vibration test specifications. Author Contributions: Conceptualization, Y.-H.P. and H.-U.O.; methodology, Y.-H.P. and H.-U.O.; software, Y.-H.P. and H.-U.O., formal analysis, Y.-H.P., S.-C.K. and H.-U.O.; validation, Y.-H.P., S.-C.K. and H.-U.O.; writing-original draft preparation, Y.-H.P.; writing-review and editing, H.-U.O.; supervision, H.-U.O.; funding acquisition K.-R.K. and H.-U.O.; All authors have read and agreed to the published version of the manuscript. Funding: This research was funded by Hanwha Systems (U-19-001). Institutional Review Board Statement: Not applicable. Informed Consent Statement: Not applicable. Aerospace 2021, 8, 201 14 of 14 Data Availability Statement: The data used to support the findings of this study are available from the corresponding author upon request. Acknowledgments: This research was supported by Hanwha Systems (U-19-001). Conflicts of Interest: The authors declare no conflict of interest. References 1. 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