Development of a High Flow-rate/High Operating ...

Development of a High Flow-rate/High Operating Frequency Nitinol MEMS Valve Myunghoon Seong1, KP Mohanchandra1, Yohan Lin2, and Gregory P. Carman1 1

Mechanical and Aerospace Engineering Department, University of California, Los Angeles 2

NASA Dryden Flight Research Center ABSTRACT

This paper presents modeling, fabrication, and testing results for a high flow-rate and high frequency Nitinol MEMS valve. ANSYS® is used to evaluate several Nitinol MEMS valve structural designs with the conclusion that a pentagonal flap with five legs produces higher frequencies and higher strengths without the inherent rotation problem present in standard designs. The Nitinol penta-leg design was fabricated using a novel bi-layer lift-off method. A PMGI polymer layer is initially used as an underlayer while a chromium layer is used as a top layer to produce a nonrotational ortho-planar Nitinol MEMS valve array. This array consists of 65 microvalves with dimensions of 1mm in circumference, 50 µm in leg width, and 8.2 µm in Nitinol thickness. Each microvalve covers an orifice of 220 µm in diameter and 500 µm in length and is capable of producing a 150 µm vertical deflection. This Nitinol MEMS valve array was tested for flow-rates in a hydraulic system as a function of applied pressure with a maximum water flow-rate of 16.44 cc/s. Keywords: MEMS, valve, Nitinol, lift-off, bi-layer, ortho-planar, high flow-rate, high frequency

1. INTRODUCTION There is a growing demand for compact hydraulic actuators in space and military applications. This requires higher actuation frequencies and larger flow-rates than currently available to satisfy power density requirement [1]. Commercially available check valves providing high flow-rates have limited frequency response. The frequency limitation (i.e. < 1 kHz) is due to the valve’s low resonance frequency producing a phase lag causing backflow and reducing performance [2, 3]. Active microvalves using piezoelectric materials or passive microvalves, on the other hand, offer higher frequency response but produce limited flow-rates (i.e. < 10 cc/s) [4, 5]. Therefore, most current valves, micro or macro, do not satisfy both high flow-rate and high frequency requirements necessary in future compact hydraulic actuators. MEMS valve structures provide high resonant frequency and, if arrays are used, make it possible to produce high flow-rate [1, 6, 7]. However, previous studies focused exclusively on either nickel or silicon materials for the valve’s structural components. These particular materials limit valve deflection because failure strains are less than 1%. In addition to material limitations, previous valve designs implemented geometries that were prone to failure. For example, spider spring geometries induce undesirable torsional stresses that cause premature failures [1, 6-10]. Therefore, new valve geometries should also be considered in addition to selecting more appropriate materials. Recently researchers began to consider novel materials such as thin film nickel titanium alloy (Nitinol) for use in thermostat valves [9]. This particular material was chosen due to its large elongation (i.e. up to 10%) along with its shape memory properties. However, manufacturing the structure using conventional MEMS processes introduces several fabrication challenges including film breakage at sharp discontinuities or edges due to wet etching issues [11]. Also, building Nitinol MEMS structures with relatively large thickness to width (i.e. > 10 %) using an isotropic etching inherently reduces mechanical strength. Gill et al. addressed some of these wet etching challenges using an ion-milling approach, however, low etch rate and low selectivity as well as limited availability virtually eliminates this approach as a

Sensors and Smart Structures Technologies for Civil, Mechanical, and Aerospace Systems 2008, edited by Masayoshi Tomizuka, Proc. of SPIE Vol. 6932, 69322F, (2008) 0277-786X/08/$18 · doi: 10.1117/12.775812 Proc. of SPIE Vol. 6932 69322F-1 2008 SPIE Digital Library -- Subscriber Archive Copy

fabrication technique for wide spread adoption [11, 12]. Therefore, thin film Nitinol as well as other new materials may require the development of acceptable fabrication processes. The Nitinol microvalve array fabricated in this paper uses a ‘bi-layer lift-off’ process to overcome fabrication issues such as undercutting at discontinuities and non-planar features associated with isotropic etching. The process takes advantage of the limited step coverage of metal deposition [13]. In the past, researchers proposed bi-layer lift-off methods for Nitinol film deposition. Nakamura et al. used chromium and polyimide bi-layers to fabricate a Nitinol loop actuator [14]. However, this fabrication method requires reactive ion etching of a polymer which is inefficient and expensive compared to classical photolithography. Roch et al. manufactured a Nitinol actuator using two different resists, however this technique is unsuitable for high temperature deposition processes (i.e. > 250 °C) required for Nitinol film [15, 16]. Consequently, a modified bi-layer lift-off method was developed in this paper. In this research, PMGI polymer layer is used as an underlayer while chromium layer is used as a top masking layer. This method permits Nitinol structures to be fabricated without introducing breakage and avoids isotropic etching issues. In addition to developing a fabrication method, a new valve geometry is introduced to achieve ortho-planar deflection without torsional stresses thereby increasing the strength of the valve. Several valve designs were considered to achieve a high flow-rate and high frequency and a finite element method (FEM) analysis was used to design and predict the performance of this Nitinol MEMS valve. Finally, this paper reports test results on a valve array for deflections and flow-rates.

2. ANALYTICAL MODEL FEM modeling was conducted using ANSYS® 12. Figure 1 shows a finite element mesh of the penta-leg orthoplannar spring structure and an illustration of the operating characteristics for the valve. The three dimensional Nitinol spring valve consists of a central flap surrounded by five legs to achieve ortho-planar deflection without rotation. The valve structure perimeter, i.e. flap and legs, is defined by a 1mm diameter circle. The central largest portion of the model, the flap, has five sides of 363 µm length and the flap moves vertically opening and closing a 220 µm diameter microchannel beneath the flap. Note that the channel diameter is not a significant factor in comparing valve performances. Each leg with 50 µm width consists of two flexible beam segments connected by a 125 µm long intermediate platform preventing in-plane rotation. The 250 µm flexible beam is anchored to a stationary base or substrate. The penta-leg design prevents out of plane rotations of the flap (i.e. rolling and pitching) while previous quad-leg designs (i.e. spider spring geometries) were unstable along diagonal axes [10]. The flap and legs in the pentaleg design are manufactured from Nitinol material in the martensite phase. Both opening and closing produce large flap deflections (δ) requiring large strains in the five legs. The large strains are accommodated with twin boundary motion (i.e. up to 10% strain) of the Nitinol in the martensite phase. When a forward pressure is applied, the flap moves vertically opening the microchannel allowing fluid to flow through the microchannel at a flow-rate Q. When a backward pressure is applied to the flap, this causes the flap to close the microchannel. This is one operating cycle and does not require external heat to induce a phase change. The resonance frequency (fr) of the spring is a function of material properties and geometry. Flexible Beam

Intermediate Platform

125 µm

250 µm

Flap

δ

Leg

Anchor

Microchannel Flow direction

Φ1mm

Figure 1. The penta-leg FEM model and operating principle.

Proc. of SPIE Vol. 6932 69322F-2

The analysis used approximately 6100 shell elements to model the response and a convergence study was performed to ensure accuracy. Each element size is approximately 10 µm × 10 µm; however, elements at corners and anchors were refined to 1 µm × 1 µm size in addition to 5 µm-radius fillet designs at corners to reduce singularities. The boundary conditions consist of anchors that are clamped and a uniform pressure applied along a circular area beneath the flap to simulate the fluid pressure produced in the microchannel. The study also included analysis of local stress variations at the anchors where failures are generally observed. In these areas, the maximum Von Mises stress values of different spring structures were compared to suggest the mechanical strength and reliability of the different structures. For comparison, quad-leg designs with the same circumference and leg widths were modeled. Using modal analysis, the FEM solution was also used to predict the fundamental resonant frequencies. The nonlinear material properties for Nitinol were obtained from tensile tests on the thin film Nitinol used in valve fabrication. The Poisson’s ratio ν = 0.3 and density ρ = 6450 kg/m3 were obtained from the literature [9]. Modeling the nonlinear twin boundary motion of Nitinol (i.e. martensite phase) was represented by a piece-wise linear curve. In addition to nonlinear material behavior, the ANSYS® analysis includes geometrically large deflections and rotations. To model both these phenomena, a multilinear elastic structural material model was used to analyze the nonlinear martensitic phase behavior of the Nitinol and large static displacement was considered in calculating the deformation of the spring structures as a function of the Nitinol film thickness and pressure. Results include deflections, strains, stresses, and resonance frequencies. The FEM predicted flap deflections at the flap’s center were used to predict flow-rates through the valve. These deflections were combined with an analytical model of an incompressible fluid flow through an orifice with an obstruction. Using the differential pressure across the orifice, the flow-rate can be estimated by Bernoulli obstruction theory [17, 18], 2 1 w/ valve flaps or

(w/o valve flaps)

(1),

4

where Q is the flow-rate, CD is the dimensionless discharge coefficient accounting for the discrepancies between incompressible steady frictionless flow and frictional flow, p is the pressure drop across the valve, ρ is the density of the fluid, D is the pipe diameter. At is the total cross sectional area of the orifice, and is a function of the microchannel diameter dch, the number of microvalves, N, and the deflection δ of the flap’s center obtained from ANSYS®. The value of CD depends on the ratio of the orifice diameter d* to the pipe diameter D, and the Reynolds number (Red*) defined by (2), where ν is the kinematic viscosity of the fluid. Once these are calculated, the appropriate CD value is obtained from plots by Johansen [18].

3. FABRICATION Figure 2 illustrates the fabrication process for the Nitinol penta-leg valve structures using a bi-layer lift-off method. The valve array is fabricated by depositing 5000 Å of sacrificial copper layer with chromium adhesion layer on a 500 µm-thick silicon substrate. An AZ4620 photoresist is then spin coated and patterned over the copper layer. The exposed copper from the MEMS valve pattern can be removed using a solution of 1(FeCl3):10(H20) followed by removing chromium with CRE-473® a chromium etchant.


Silicon

Cr/Cu evaporation and pattern

PMGI resist coating and pattern Cr mask layer

PMGI pattern and Nitinol sputtering

Lift-off and Nitinol annealing

Backside photolithography

DRIE back Channel and release

Silicon

Chromium Copper

PR

PMGI

Nitinol

Figure 2. Fabrication process flow of Nitinol MEMS valve.

A thick SF-11 PMGI layer is coated using multiple spin processes and 1500 Å of chromium layer is subsequently deposited. The PMGI is spin coated at 500 rpm (3.4 µm thick) and baked at 195°C for 2 minutes. Repeating the PMGI coating 3 ~ 4 times makes it possible to achieve a thick PMGI layer over 10µm. An AZ5214 photoresist is used to pattern the chromium with the chromium etchant. The patterned chromium layer represents the masking layer of the PMGI layer which is patterned by deep UV exposure and XP 101A PMGI developer. After PMGI development, O2 plasma etching is used to minimize the polymer residues. The Nitinol is deposited on the front side of the substrate in a DC magnetron sputtering system. The target used for the sputtering system has a composition of 52 at.% titanium and 48 at.% nickel. The target-substrate distance is kept at 4cm while holding the Argon pressure at 1.4x10-3 Torr. A substrate holder translates horizontally to achieve a better film uniformity. After deposition of the Nitinol film, lift-off process is followed to remove unnecessary Nitinol film deposited on top of the chromium and PMGI bilayer. This lift-off process produces patterns for the MEMS valve flaps and tethers without etching of the Nitinol film. After the lift-off step, Nitinol film is crystallized by annealing at 540 oC for 30 minutes under vacuum at better than 10-6 Torr. Ho et al. described the sputter deposition and annealing processes in detail [16]. The backside of the substrate is spin coated and developed with NR5-8000 negative photoresist and RD6 developer to produce a pattern necessary for channel fabrication. The patterns are an array of circles, each circle having a diameter of 220 µm. The silicon is then etched through the substrate with DRIE. The final release step is followed by the removal of the sacrificial layers.

4. TEST SETUP Nitinol film is characterized using a stress-strain test. An MTS Tytron Microforce Actuator 250®, a mechanical testing machine, is used to measure stress-strain behavior of the film. The stress-strain behavior is examined at room temperature (i.e. Young’s Modulus and the stress when detwinning begins in martensite phase) using a piece of Nitinol film with dimension of 21.2 mm length, 3 mm width and 5 µm thick placed in the sample holder of the MTS Tytron machine. The sample is loaded in force control mode with a rate of 0.05 N/s. The tensile load is increased until the sample strains about 3.5 % and then it is decreased to 0 at a rate of 0.05 N/s.


An Omega stereo microscope with 45x magnifications and a fiber optical illuminator and a Nikon Coolpix 950 camera are used to observe the microvalve structure operation. The microscope allows observations of the 3D movement in the structure; therefore, the ortho-planar movement as well as the rotation of the micro-structure can be observed. A microvalve array is placed inside a fixture connected to a pneumatic reservoir applying constant pressures ranging between 0 to 137.9 kPa at the backside of the valve array. In addition to the microscopic setup, a Keyence LK-081 laser displacement measurement (LDM) device with a resolution of 3 µm is used to measure the microvalve’s failure deflection. The flow-rate of deionized (DI) water (ρ=998 kg/m3 and ν=1.005×10-6 m2/s at 20oC) through the microvalve array was measured. A constant pressure ranging from 0 to 103.5 kPa incremented in steps of 6.9 kPa (e.g. 6.9, 13.8, etc.) is applied to the valve using DI water. The valve array is connected to a Φ4.8 mm (D) pipe supplying the DI water. The flow-rate at each constant pressure condition is determined by measuring the weight of DI water flowing through the array over a specific time (e.g. 20 seconds for high flow-rates and 60 seconds for low flow-rates). Leakage is measured by applying flow in the backward direction. Finally, the flaps of the microvalves were removed to measure the flowrate thorough an orifice without an obstruction.

5. RESULTS Figure 3 shows the measured stress-strain behavior of the Nitinol film in the martensite phase measured at room temperature. The experimental data can be approximated by two linear stages in the loading curve. The first linear stage corresponds to elastic deformations up to 0.7% with a Young’s modulus of 19.7 GPa. Above 0.65 % strain, inelastic deformations associated with twin boundary motion accommodate strains up to 3.7 %. In the detwinning region the secant modulus is 6.3 GPa. During unloading, a non-linear decrease in the stress-strain behavior is observed, which is attributed to twinning due to residual stresses, and when the load becomes 0, a residual strain of 2.5 % is observed. The residual strain is recovered by applying a stress in the opposite direction and again moving twin boundaries. This complete cycle, while hysteretic, is reversible without damaging the material as contrasted with dislocation motion. 350 EXP Curve Fit 300

Stress [MPa]

250

200

150

100

50

0

0

0.005

0.01

0.015

0.02 Strain

0.025

0.03

0.035

0.04

Figure 3. The stress-strain behavior of the Nitinol film.

ANSYS® finite element Von Mises stress contour plots for a 10 µm thick quad-leg and the penta-leg designs subjected to 69 kPa pressure through a channel of 220 µm diameter are presented in Figure 4. The top two figures (figures 4a and 4b) show the stress distribution in the quad-leg and penta-leg designs and the bottom two figures (figures 4c and 4d) show the stress distribution on the flap above the microchannel for each design. The stress distributions on the flap illustrate the principal differences in the two designs. For the quad-leg, the stress is highest at the outer rim and decreases in the radial direction toward the flap’s center. For the penta-leg, however, the stress is highest at the flap’s center and decreases in the radial direction toward each leg. This stress distribution difference is attributed to warpings (i.e. rotation) present in the quad-leg design. In addition to this difference, one can see the asymmetric stress


distribution about plane A-A’ (see figures 4a) for the quad-leg design, while the penta-leg shows approximately a symmetric stress distribution about a similar plane B-B’ (see figures 4b). In the top two figures, a superposition of deformed and undeformed images of the quad-leg and the penta-leg configurations are also presented. These superpositions clearly show a significant flap rotation in the quad-leg design whereas there is an absence of rotation in the penta-leg design. This rotation is attributed to the asymmetric stress distribution in the quad-leg design, while symmetric stress distribution of the penta-leg design balances out rotation of the each flexible beam.

(b)

A

(a)

A’

Rotation

No rotation (d)

(c)

Figure 4. (a) Von Mises stress contour plot of the quad-leg design. (b) Stress contour plot of the penta-leg design. (c) Stress distribution on the flap of the quad-leg design. (d) Stress distribution on the flap of the penta-leg design.

Figure 5 shows plots of the resonance frequencies and the maximum Von Mises stresses at the anchors as a function of Nitinol film thickness ranging from 7 µm to 15 µm. For the maximum stress calculation, a 69 kPa pressure is applied to a 220 µm diameter channel. The solid line with triangles represents the variation of the 1st resonance frequency of the quad-leg design as a function of the film thickness, and dashed line with inverted triangles is for the penta-leg design. The penta-leg design shows substantially higher resonance frequencies than the quad-leg design at a given film thickness and increases more rapidly for increasing film thickness. In figure 5b), the penta-leg design shows lower stress concentration at anchors than the quad-leg design. This value is larger than what would simply be attributed to the addition of another leg, i.e. 25% decrease. These results show that the penta-leg design can achieve better responsiveness (i.e. higher resonance frequency), and reduce structural failures compared to the quad-leg design. 22

300

σ anchor (quad-leg)

Fres (quad-leg) Fres (penta-leg)

20

σ anchor (penta-leg) 250

Maximum Stress at Anchors [MPa]

Resonant Frequency (kHz)

18

16

14

12

10

8

200

150

100

6

4

7

8

9

10

11 12 Film Thickness [ µm]

13

(a) Resonance frequency vs. film thickness

Figure 5.

14

15

50

7

8

9

10

11 12 Film Thickness [ µm]

13

14

15

(b) Maximum stress at anchor vs. film thickness

Comparison of the quad-leg and the penta-leg designs in resonance frequencies and stresses.


(a)

(b)

200µm

(d)

(c)

Figure 6.

(a) Nitinol MEMS valve array. (b) SEM graph of a single valve. (c) Closed valve. (d) Deflected valve

Several pictures of an 8.2 µm thick Nitinol penta-leg ortho-planar valve array using the new bi-layer lift-off method is shown in Figure 6. Figure 6a shows an array of 65 microvalves (N) superpositioned over a quarter for reference purposes. This valve array was used in experimental measurement of flow-rate. Each microvalve covers a microchannel of 220 µm in diameter (dch) and 500µm in length. Figure 6b shows a scanning electron microscope (SEM) graph of a single valve and no issues inherent for wet-etching were found using the bi-layer lift-off method. Figures 6c and 6d show photographs of an operating valve. Figure 6c shows a closed valve when there is no pressure applied, and figure 6d shows an open valve when a static pneumatic pressure is applied. Due to the depth of focus limitations of the microscope setup, the pictures are not taken at the same angles. The important feature in figures 6c and 6d is that penta-leg deflections occur without rotation. Valve failure was only observed at displacements greater than 150 µm flap displacement. 25 with flaps (EXP) with flaps (MODEL) w/o flaps (EXP) w/o flaps (MODEL) Backpressure leakage

Water Flow rate at room temperature [cc/s]

20

15

10

5

0 0

Figure 7.

20

40

60 Pressure [kPa]

80

100

120

Analytical and measured flow-rates as a function of pressure applied.

Figure 7 provides plots of flow-rate as a function of applied pressure from 6.9 kPa to 103.5 kPa for the valve array shown in Figure 6a. The five data sets in the figure represent experimental (data points) and analytical (solid and dash-


dot lines) flow-rates with microvalve flaps or without flaps, and flow leakage when backpressure is applied. Referring to the experimental flow-rate with flaps (i.e. squares in the figure), the flow-rate increases linearly with pressure and exceeds 10 cc/s at approximately 65 kPa. The 10 cc/s represents a typical operating requirement for the valves. On the other hand, experimental data with flaps removed (i.e. circles) shows an increase proportional to the square root of pressure. Comparing these data with the analytical model (equation (1)) required the Reynolds numbers and the discharge coefficient CD values. The calculated square root values of the Reynolds numbers (equation (2)) for the experimental flow-rates with flaps range between 51 to 106, while the ratios of the orifice diameter d* (equation (1)) to the pipe diameter D calculated using FEM deflection results are between 0.12 to 0.39. Similarly, the Reynolds numbers for the flow-rates without flaps were calculated using experimental flow-rates while the ratio d*/D fixed (i.e. 0.37) to determine CD values. Experimental flow-rate data with flaps removed is in good agreement with the analytical model (i.e. the dash-dot line). However, the model (i.e. the solid line) fails to predict the flow-rate with the flaps. At lower pressure (i.e. < 30 kPa), the flow-rate is in good agreement with the model; however, the discrepancy becomes larger and larger as the pressure increases and the flow-rate tends to follow the flow-rate without flaps. Therefore, it can be determined that the flow-rates with microvalve flaps are affected by the valve displacements at low pressure while the channel geometry becomes more dominant factor at high pressure. Finally, the valve array produced no measurable backpressure leakage (i.e. triangles) up to 137.9 kPa.

6. CONCLUSION A Nitinol MEMS valve was fabricated using a bi-layer lift-off method. The new method shows better results especially at discontinuities (i.e. no undercut, film brakeage) than previous wet-etching fabrication methods. ANSYS® simulations as well as experiments proved that a new penta-leg spring design improves mechanical strength and reliability performance over the quad-leg spring design. For the fabricated Nitinol valve, it is expected to have a resonant frequency of 11.9 kHz in air using FEM. Flow-rates over 10 cc/s can be achieved with DI water.

ACKNOWLEDGEMENT This research was supported by NASA Dryden Flight Research Center’s Compact Piezoelectric Hydraulic Pump Program managed by Yohan Lin.

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