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An effective thermal conductance tuning mechanism for uncooled microbolometers
Infrared Physics & Technology 57 (2013) 81–88
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Infrared Physics & Technology journal homepage: www.elsevier.com/locate/infrared
An effective thermal conductance tuning mechanism for uncooled microbolometers Nezih Topaloglu a,⇑, Patricia M. Nieva b, Mustafa Yavuz b, Jan P. Huissoon b a b
Yeditepe University, Department of Mechanical Engineering, 34755 Istanbul, Turkey University of Waterloo, Department of Mechanical and Mechatronics Engineering, 200 University Avenue West, Waterloo, ON, Canada N2L 3G1
h i g h l i g h t s " A thermal conductance tuning mechanism for uncooled microbolometers is proposed. " A pixel-by-pixel tuning is realized and a stopper mechanism is used. " Thermal conductance can be tuned by a factor of three. " The tunability is highly linear.
a r t i c l e
i n f o
Article history: Received 28 May 2012 Available online 3 January 2013 Keywords: MEMS Uncooled microbolometers Thermal modeling Tunable thermal conductance
a b s t r a c t With the increasing demand on infrared (IR) detectors for imaging harsh environment processes, widening the application range of uncooled microbolometer arrays has become an important research area. An efficient way of increasing this range is tuning the thermal conductance of the microbolometer array using electrostatic actuation, which is usually achieved by directly applying an actuation voltage to the substrate. However, this method does not allow pixel-by-pixel actuation, limiting the tunability. In this paper, we present a new method of actuation which uses the micromirror located below the microbolometer as the actuation terminal. We demonstrate that using micromirror actuation, the thermal conductance can be tuned by a factor of three. An analytical model to calculate the thermal conductance of this new type of microbolometer is presented. Results of the model are compared to finite element simulations and experimental measurements on a test structure fabricated for this purpose, showing good agreement. The new tuning mechanism provides a fairly linear thermal conductance tunability, thus making it a promising thermal conductance controlling mechanism for adaptive IR detectors. Ó 2012 Elsevier B.V. All rights reserved.
1. Introduction The demand for uncooled microbolometer focal planar arrays (FPAs) has increased significantly over the last few years, due to their high sensitivity and low cost. Due to the recent growth in the application areas of thermal imagers, such as high temperature process imaging [1], increasing the device dynamic range has become one of the main challenges in microbolometer research. For instance, when a high temperature process (e.g. 3000 K during an exothermic reaction) is imaged with a microbolometer FPA, the high radiation power can increase the temperature of the device by about 100 °C [2]. This increase can lead to a number of undesirable effects, such as nonlinear response, plastic deformation of both the arms and the microplate, and ultimately device failure. To overcome these problems, the thermal conductance ⇑ Corresponding author. E-mail addresses: [email protected] (N. Topaloglu), pnieva@ uwaterloo.ca (P.M. Nieva). URL: http://simslab.uwaterloo.ca (P.M. Nieva). 1350-4495/$ - see front matter Ó 2012 Elsevier B.V. All rights reserved. http://dx.doi.org/10.1016/j.infrared.2012.12.039
(G) can be varied directly by contacting the detector with its substrate using electrostatic actuation [3,4], where the thermal conductance of the microplate is controlled by the thermal contact conductance (TCC) of the detector-substrate interface [4]. However, thermal conductance is a parameter that directly affects the responsivity and the response time of a microbolometer. As the thermal conductance decreases, both the responsivity and the response time increase. Hence, any variation in thermal conductance results in a trade-off between the responsivity and response speed. Therefore, a microbolometer FPA with adjustable thermal conductance can give the user the opportunity to switch from a high temperature resolution and low frame rate mode to a low resolution and high frame rate mode, depending on the application. Previous thermal conductance tuning mechanisms apply a certain voltage directly to the substrate to actuate the microplate [3,4]. Nonetheless, this type of actuation does not allow pixel-bypixel thermal conductance tuning. In addition, due to the large contact area at pull-in, the change in thermal conductance is sudden and very high, which prevents linear thermal conductance tunability.
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Previous research work in our group presented a new microbolometer design that used a stopper mechanism to adjust the thermal conductance [5]. The stopper mechanism allowed for a fairly linear pixel-by-pixel thermal conductance tunability due to the small contact area at pull-in. In this same work, it was demonstrated that the thermal conductance for a 120 120 lm2 microplate size microbolometer was doubled, with respect to the unactuated state, when an actuation voltage of 12 V was used. In this paper, we propose an improved microbolometer design which also includes a stopper mechanism. The microplate size of the improved microbolometer is 80 80 lm2 which is considerably smaller when compared to the 120 120 lm2. This size is also closer to the sizes of commercial microbolometers which range from 17 17 lm2 to around 40 40 lm2 [6]. In addition, the arm length is increased from 160 lm to 180 lm, for increasing the thermal isolation and to decrease the contact voltage. The design and fabrication of the proposed tunable thermal conductance mechanism are presented in Section 2, and then compared to existing mechanisms. Section 3 includes an analytical thermal conductance model, which calculates the thermal conductance of the device in terms of the actuation voltage. A comparison of the analytical results, with finite element simulations, electrical capacitance measurements and thermal conductance measurements is presented in Section 4. Finally, the conclusions are drawn in Section 5. 2. Design and fabrication Fig. 1 shows a schematic of a tunable thermal conductance mechanism for microbolometers, demonstrated by Song and Talghader in 2002 [4]. In this design, the microplate is actuated by applying a bias voltage to the substrate (Fig. 1b). In this case, the authors demonstrated that thermal conductance can be increased up to four times when the detector snaps down completely onto the substrate. Later, the same authors demonstrated that the tunability range of the thermal conductance can be increased with a similar improved design, which used the same substrate biasing method [2]. Although biasing the substrate adds great tunability to the thermal conductance, it has a disadvantage. The bias voltage applied to the substrate results in the actuation of the whole microbolometer pixel array, which makes pixel-by-pixel thermal conductance tuning difficult. This problem could be avoided by biasing the
Fig. 1. Cross-sectional schematic of the microbolometer thermal conductance tuning mechanism demonstrated in [4] in (a) unactuated state and (b) actuated state. (Not drawn to scale.)
micromirror plate, which is located below the IR sensitive microplate, instead of the substrate. Since the micromirror is used only for increasing the infrared absorption of the microbolometer pixel; the effect on the normal operation of the microbolometer when a voltage is applied to it will most likely be negligible. However, to maintain a continuous electrostatic actuation force and safe operation, there should not be any electrical contact between the microplate and the micromirror. This can be achieved by preventing the snap-down of the microplate onto the micromirror once a voltage higher than the pull-in voltage is applied. Depositing a sufficiently thick non-conductive material on the microplate that can electrically isolate the micromirror could be a potential solution. However, in this case, the thermal conductance will reach its maximum value abruptly, resulting in a digitized one-step thermal conductance tuning. To precisely control the thermal conductance according to the environmental conditions, a thermal conductance that changes linearly with the applied voltage is required. To overcome the problems mentioned above, a stopper mechanism was designed and it is illustrated in Fig. 2. In this design, the contact between the microplate and the micromirror is avoided using stoppers. Tuning the thermal conductance is achieved using electrostatic actuation of the stoppers. The stoppers are rigid structures anchored to the substrate, and they extend into the gap between the microplate and the base plate. During operation, as the microplate deflects sufficiently towards the biased micromirror, the stoppers prevent the upper plate from deflecting further upon contact, thus avoiding the contact with the micromirror (see Fig. 2b). At the same time, the thermal conductance increases, since the heat flows to the substrate mainly through the stoppers, which provide a good heat link between the microbolometer and the substrate. When the bias voltage is decreased to 0 V, the microplate reverts back to its original position, thus eliminating the heat link provided by the stoppers (Fig. 2a). In that case, thermal conductance will decrease back to its lowest value, which is then only limited by heat conduction through the microbolometer arms. To demonstrate this stopper mechanism, several test structures were fabricated using PolyMUMPs (Polysilicon Multi-User MEMS Process) [7]. PolyMUMPs is a three-layer polysilicon process, which has well-defined process steps and design rules. Fig. 3 shows the SEM image of a fabricated test structure. Since the stopper should extend in between the microplate and the base plate (mirror), the base plate, the stoppers and the microbolometer were made by the Poly 0, Poly 1 and Poly 2 layers, respectively. The air gap between the microplate and micromirror is 2.75 lm. It should be noted that in an actual microbolometer, the micromirror is made of metal.
Fig. 2. Cross-sectional schematic of the microbolometer with the proposed stopper, in (a) unactuated state and (b) actuated state. (Not drawn to scale.)
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At the unactuated state, the stoppers do not have any effect on the thermal conductance, and the conduction through the microbolometer arms is the dominant heat conduction mechanism. Therefore, the thermal conductance can be calculated by applying Fourier’s law of conduction to the arms.
Gunact ¼
2kPolySi wPolySi t PolySi 2kgold wgold tgold þ ; LPolySi Lgold
ð1Þ
where k, w, t and L are the thermal conductivity, width, thickness, and length of the corresponding arm layer, respectively. At the actuated state, the microplate contacts the stoppers, forming a new heat link between the microplate and the substrate. The temperature decrease at this heat link is mainly determined by the thermal contact conductance between the microplate and the stoppers. Therefore, the thermal conductance at the actuated state (Gact) can be stated as
Gact ¼ Gunact þ Gc ;
Fig. 3. Scanning electron microscope (SEM) image of the test structure used in this paper: a 80 lm 80 lm microplate, connected to the substrate with two 180 lm 10 lm arms.
However, for these test structures, the Poly 0 layer is used as the base plate of the electrostatic actuator instead of a metallic layer, due to the limitations of the PolyMUMPs process. Nevertheless, this situation is not expected to affect the experimental results, since the micromirror material has no effect on the electrostatic actuation. The authors would also like to acknowledge here that the addition of stoppers results in added complexity and a higher pixel size, which might not be an option for some FPA applications. However, a compromised solution could result from further optimization of the shape, materials and dimensions of the microbolometer pixel, specially for high temperature applications such as the ones mentioned before. This however is not within the scope of this paper. The area of the microplate of the test structure used for the purpose of this paper is 80 80 lm2. This area is considerably smaller than the microplate area of the test structure presented in Ref. [5], which showed a similar performance. It should be noted that stateof-art microbolometer designs can have microplate areas as small as 17 17 lm2 [6], thus reducing the size of a microplate while maintaining the performance of new microbolometer designs could be used as a measure of improvement. To improve and prevent stiction during final release, 7 lm 7 lm holes and 4 lm 4 lm dimples are used. The microbolometer arms are 180 lm long and 10 lm wide and were coated with the Au metallization, to make sure that the resistive heating occurs mainly within the microplate. 3. Thermal conductance modeling In this section, modeling of the thermal conductance of the proposed mechanism is presented. The main heat transfer mechanisms involved in the operation of a microbolometer are convection and radiation between the pixel and its surroundings, and conduction to the substrate. For thermal isolation, microbolometers are packaged in vacuum, which means convection can be neglected. Radiation can also be assumed to be negligible, since the pixel never reaches extremely high temperatures during operation [8]. Therefore, for both the unactuated and the actuated state, the thermal conductance models, for our particular case, are based on the conduction of the heat to the substrate through the arms and the stoppers.
ð2Þ
where Gc is the thermal contact conductance between the microplate and the stoppers. The term Gc stems from the fact that the contacting surfaces are not perfectly smooth and flat. Therefore, as the two surfaces contact, they will touch only at a portion of the apparent contact area, forming microgaps between them. As a result, the heat is transferred only through the contacting asperities, which results in a temperature drop across the contact. A small portion of a contact between two nominally flat surfaces (also called conforming rough surfaces) is shown schematically in Fig. 4a. Each surface has a mean plane, which is shown in the figure with a dotted-dashed line. The surface roughness and the absolute mean asperity slope of the surface i are defined by ri and mi, respectively, both of which can be determined from the corresponding surface profile data as [9]
ri
sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Z 1 L 2 ¼ y ðxÞdx; L 0 i
mi ¼
1 L
Z 0
L
dyi dx; dx
ð3Þ
ð4Þ
where yi(x) is the distance of the points on surface i from the mean plane, and L is the length of a trace. The contact between the rough surfaces can be transformed into the contact between a flat surface and a rough surface, where the effective roughness, r, and the effective absolute mean asperity slope, m, are defined as [9]
r¼
qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r21 þ r22 ;
ð5Þ
m¼
qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi m21 þ m22 :
ð6Þ
A small portion of the contact after this transformation is shown in Fig. 4b. Gc depends on many parameters, such as the surface roughness and the contact pressure. Various models are available for determining Gc [9]. Most of these models are based on the assumption that surface asperities are distributed randomly over the contact area, and they have Gaussian height distribution about some mean plane passing through each of them. These models are called fully Gaussian (FG) models. However, it was reported that at low contact pressures, the fully Gaussian models systematically underestimate the thermal contact conductance [10]. An explanation to this inaccuracy at low contact pressures is proposed by Milanez et al. [11]. There, it was reported that the real surfaces may have Gaussian surface distribution up to around 4.5r, however they generally do not have asperities whose height is above 4.5r [11]. Therefore,
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Fig. 4. Schematic showing a small portion of (a) a contact formed between two conforming rough surfaces, (b) the same contact transformed into the contact of a flat and a rough surface. Reprinted from [9].
the distance between the mean planes at low contact pressures is shorter than expected and the contact conductance is greater than expected. As the contact pressure increases, more number of shorter asperities come into contact, which explains the accuracy of the fully Gaussian model at higher pressures. As a solution to this inaccuracy, a truncated gaussian (TG) model was developed [11,12]. This model assumes that the heights of the surface asperities follow the Gaussian distribution up to a defined value of ztrr, where ztr is the relative truncation level. Due to its higher accuracy in low contact pressures, the TG model is employed for determining Gc. Assuming the asperities deform elastically upon contact, Gc can be expressed as [12]
sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi m 1 ; Gc ¼ 1:25Aa ks ðP=He Þ0:95 ½1 þ 1=f 0:9289 1 1þf r
ð7Þ
where
pffiffiffiffiffiffiffi f ¼ ðP=He Þ 2p ztr exp z2tr =2 :
ð8Þ
In Eqs. (7) and (8), Aa is the apparent contact area, P is the contact pressure and He is the effective elastic microhardness, given by
m E ; He ¼ pffiffiffi 2 2 2 1m
ð9Þ
where E and m are the Young’s modulus and the Poisson’s ratio of polysilicon, respectively. To find P, the net force Fnet acting on the microplate upon contact should be determined. Approximating the microbolometer to a moving plate capacitor, where the moving plate is the microplate, Fnet is given by
F net ¼ keq ðg 0 gÞ
0 AV 2act 2g 2
area, P can be found. The contact voltage Vcontact, which will be defined here as the voltage where the microplate touches the stoppers, also needs to be found, to determine the beginning of the actuated state. Since the height of the stopper is smaller than the height at which pull-in occurs, Vcontact can be found by equating Fnet to zero, and g to 2g0/3, which yields
V contact
sffiffiffiffiffiffiffiffiffiffiffiffiffiffi 8keq g 30 ¼ : 270 A
ð11Þ
4. Experiments and discussion 4.1. Optical profiler measurements A Wyco NT100 optical profiler was used to determine the thicknesses of the air gap and the polysilicon films forming the test pixel shown in Fig. 3 after fabrication. They were also used to determine the effective surface roughness parameters (r and m), and the surface truncation parameters (ztr). The parameters, extracted from the optical profiler data are tabulated in Table 1. The optical profiler data shows that the gap between the microplate and the base plate is 2.7 lm at the unactuated state and 1.7 lm at the actuated state. The gap thickness of 1.7 lm is less than the design value of the Oxide 1 layer (2 lm [7]), which shows
Table 1 The thicknesses and roughness parameters extracted for the test structure shown in Fig. 3 using data from an optical profiler. Thicknesses (lm)
;
ð10Þ
where A is the microplate area, g0 is the initial gap thickness, and g is the final gap thickness, 0 is the permittivity of free space, and keq is the equivalent spring constant of the arms, which can be found by beam bending analysis. Dividing Fnet to Ac, which is the total contact
Poly 0 Poly 1 Poly 2 Gold Initial gap (g0) Gap after contact (g)
Roughness parameters 0.6 2.0 1.3 0.5 2.7 1.7
r1 m1
r2 m2 ztr
12.14 nm 0.042 2.06 nm [13] 0.007 2.44
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that the microplate successfully contacts the stoppers at the actuated state. In Figs. 5 and 6, the 3D surface profile and X and Y profile of the test structure captured by WYKO NT100 Optical Profiler is plotted, when the actuation voltage is 12 V. In a parallel plate capacitor, pull-in occurs when the gap between the capacitor plates is less than two-thirds of the initial gap. In our case, the results of the optical profiler measurements show that the gap after pull-in (g) is less than two-thirds of the initial gap (g0) value. Hence, Eq. (11) can be used as an analytical expression to calculate the contact voltage. To calculate r and m, the surface roughness of the contacting surfaces should be determined. Taking the top surface of the stoppers (Poly 1) as surface 1, r1 and m1 is calculated from the optical profiler data as r1 = 12.14 nm and m1 = 0.042. The surface roughness parameters of the other contacting surface (r2 and m2), which is the bottom surface of the microplate (Poly 2), cannot be measured by the optical profiler. Instead, r2 is taken from [13] as 2.06 nm and m2 is estimated according to the formula m2 = m1(r2/r1). Using these values, r and m can be found (r = 12.31 nm, m = 0.043). The truncation level ztr is determined by extracting the height of the highest asperity from the surface profile data and dividing it to the surface roughness, which yields ztr = 2.44. 4.2. Finite element model simulations To determine the contact voltages of the test structures in the design stage, a finite element model (FEM) was constructed using the commercial finite element analysis (FEA) software ANSYSTM. The total number of elements of the constructed model was around 100000. The structure was modeled with the 3-D 20 node structural SOLID 95 element, whereas the air surrounding the structure was modeled with the electrostatic SOLID 122 element. The mechanical properties of polysilicon and gold were taken from the PolyMUMPs manual [7]. For simplicity, the stoppers were not included in the FEM. After the construction of the model, a coupled electrostatic-structural solution was performed using the ESSOLV macro. The electrostatic solver calculates the mechanical forces based on the potential difference, and the structural solver uses the force distribution to calculate the deflection. If convergence is not reached after the first solution step, the deflection data is used again by the electrostatic solver to recalculate the force distribution. This loop continues until convergence is reached. Using the gap and thin-film thicknesses obtained from the optical profiler data (Table 1), the contact voltage calculated using the FEM model was 13.5 V. The theoretical contact voltage, i.e. the contact voltage calculated by Eq. (11), is found as 12.4 V, which is fairly close to the contact voltage found by the FEM. The small discrepancy between these values are mainly attributed to the limitations introduced in the theoretical model when assuming a perfectly flat microplate and when calculating the equivalent spring constant of the arms using Euler Bernoulli beam theory.
Fig. 5. 3D image of the test structure created by WYKO NT100 Optical Profiler.
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4.3. Electrical capacitance measurements A reliable way to measure the contact voltage is to measure the electrical capacitance. As the actuation voltage increases, the electrical capacitance increases as the result of the decrease in gap (g). A sudden increase is expected once contact voltage is reached. The voltage versus capacitance plot for the device under study, measured with an Agilent E4980A Precision LCR Meter is shown in Fig. 7. As can be seen from Fig. 7, a jump in the capacitance is observed roughly at 11 V, which is the voltage at which the microplate touches the stoppers. The increase in capacitance after 12 V is less than 0.2%, which for the effects of this paper can be assumed as negligible. The authors believe that compared to the FEM (13.5 V) and theoretically calculated contact voltage (12.4 V), the lower experimental contact voltage is mostly due to the effect of the adhesion force between the stoppers and the microplate, which was not considered in the FEM and theoretical contact voltage calculation. 4.4. Thermal conductance measurements The thermal conductance was measured under vacuum (<4 mTorr) using the Joule heating method [14]. In this method, heat is generated by biasing the microplate with a voltage or current source under vacuum. By measuring the current at each voltage step, a plot of dissipated power versus resistance is obtained. G is calculated from the slope of this curve, according to the following formula [15]
R ¼ R0 þ
1 R0 aI2 R; G
ð12Þ
where I is the current, R0 is the resistance at 300 K and a is the temperature coefficient of resistance (TCR), found as 5.2 104 K1 by obtaining the temperature versus resistance plot in a temperature controlled environment. The power versus resistance measured for the microbolometer pixel shown in Fig. 3, is plotted in Fig. 8. In the unactuated state, G is found to be 1.29 105 W/K. This value is in good agreement with 1.30 105 W/K, which is the thermal conductance calculated by Eq. (1). With an actuation voltage of 11 V, the thermal conductance increases to 2.44 105 W/K. As discussed previously, this increase in G is expected, due to the new heat links formed by the contact of microplate to the stoppers. As the actuation voltage is further increased, the pressure of the microplate on the stoppers increases, resulting in higher TCC. This explains the increase in thermal conductance to 3.71 105 W/K, with an actuation voltage of 35 V. After around 35 V, the devices failed mostly due to overheating, as a result of breakdown failure. It should be noted that due to the limitations of the PolyMUMPs process, the dimensions of the fabricated structures are not ideal (e.g. arms cannot be as thin and materials cannot be changed), so thermal conductances and contact voltages are not low enough for it to be used as a commercial microbolometer [16]. By using a custom fabrication and state-of-the-art fabrication equipment, lower actuation voltages and higher tunability in thermal conductance are expected. The tunability range of the thermal conductance can be found by dividing the thermal conductance at its highest value as a result of applied voltage, to thermal conductance at the unactuated state. In Fig. 9, the measured thermal conductance is compared to calculated values using Eq. (2) of the thermal conductance model presented in the previous Section. From Fig. 9, it can be concluded that the results obtained using the analytical model shows a similar trend when compared to the ones obtained experimentally. However, it underestimates the thermal conductance at the
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Fig. 6. Profile of the test structure in (a) X and (b) Y directions.
Fig. 7. C–V response of the test structure shown in Fig. 3 measured by an LCR meter. Note that there is a sudden jump in the capacitance roughly at 11 V.
Fig. 8. A plot of the power versus resistance, for the test structure with stoppers shown in Fig. 3. The top curve corresponds to the unactuated state (Vact = 0).
actuated state. In addition, due to the discrepancy in the experimental and theoretical contact voltage, the onset of the actuated state of two curves differ by around 2 V. The inaccuracy in the contact voltage estimation also leads to an inaccuracy in contact pressure estimation, which we can be attributed as the main reason of the underestimation of thermal conductance.
To observe the effect of stoppers on the contact voltage and the thermal conductance, Fig. 10 shows the thermal conductance (G) versus actuation voltage of two different microbolometer structures. One of these structures is the microbolometer in Fig. 3 (structure A). The second structure is a microbolometer that has the same geometry and has been fabricated in the same batch
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Fig. 9. Thermal conductance versus actuation voltage of the test structure. The solid curve corresponds to the experimental data and the dashed curve corresponds to the analytical thermal contact model.
Fig. 10. Thermal conductance versus actuation voltage of test structures with and without the stopper layer.
using PolyMUMPs. This device still keeps the polysilicon micromirror at the bottom but does not have the stoppers (structure B). To avoid the electrical contact between the actuation terminals, structure B is actuated by biasing the substrate, which is isolated from the surface structures by a 600 nm silicon nitride layer. As it can be observed, before contact, both structures exhibit a constant thermal conductance (1.29 105 W/K for structure A and 1.25 105 W/K for structure B). At contact voltage, the thermal conductance of structure B jumps from 1.25 105 W/K to 9.51 105 W/K, which corresponds to a tunability range of around 8. On the other hand, the thermal conductance of structure A at contact becomes 2.44 105 W/K. This is expected since at contact, the microplate in structure B snaps down to the substrate, resulting in a higher contact area than that on structure A at contact. In addition, when the structure B is at contact, the gap between the microplate and the substrate is smaller (1.1 lm corresponding to the nitride and micromirror layers, refer to Fig. 2), compared to the contact of structure A, resulting in a higher electrostatic attraction. From Fig. 10, we can also see that for the test structure B, as the actuation voltage is increased from 0 to the contact voltage, there is less than a 0.3% change in G. Therefore, the thermal conductance can only be tuned to two different values, the one before, and the one after contact. In comparison, the G of the structure with stoppers increases linearly at a rate of 5 105 W/K/V, due to the increase in TCC between the stoppers and the microplate. This linear response provides us with a continuous tunability range as well as the ability to control the thermal conductance more precisely. The tunability range of test structure B is calculated to be around 3. From Fig. 10, it should be also inferred that the contact voltage of structure B is measured to be higher than that of structure A, which shows the effect of stoppers-microplate adhesion force on the contact voltage. As it was mentioned before, a common problem that may arise during electrostatic actuation is the in-use stiction. Although the
devices type A are expected to exhibit low adhesion forces due to their smaller contact area, we observed in-use stiction during testing, when the voltage reaches contact voltage. On the other hand, no evidence of in-use stiction was observed in the structures type B. This may be attributed to the fact that the stoppers limited the deflection of the microplate, which resulted in a restoring force smaller than the adhesion force, when the actuation voltage was removed. We believe that in order to avoid stiction, the gap between the stoppers and the microplate should be higher. This way, the microplate will need to deflect more before contacting the stoppers, which in turn will increase the restoring force. An adequate increase in the restoring force will easily bring the microplate back to its original position when the actuation voltage is removed. This can be achieved if the stoppers are fabricated closer to the base plate by using a customized fabrication process instead of PolyMUMPs. In addition, since the stopper material affects the adhesion force, choosing an appropriate stopper material could further decrease this force. The problems encountered with inuse stiction during experimental testing prevented us from conducting repeatability tests using the same device. However, onset actuation voltages measured in identical test structures as the one shown in Fig. 3, from the same batch but different chips, were also around 11 V.
5. Conclusion An efficient thermal conductance tuning mechanism for uncooled microbolometer, based on electrostatic actuation has been proposed. Instead of biasing the substrate and to realize a pixelby-pixel tuning, a stopper mechanism is used to prevent electrical contact between the electrostatic actuation terminals. We demonstrate that this mechanism linearizes the thermal conductance tuning after the onset actuation voltage needed to contact the microplate with the stoppers. A thermal contact conductance
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model was used for estimating the thermal conductance as a unction of the actuation voltage. Using this stopper mechanism, the thermal conductance of the microbolometer was increased with a linear tunability ranging from 2.44 105 W/K to 3.71 105 W/K. Considering the lowest and highest thermal conductance measured (1.29 105 W/K and 3.71 105 W/K), we observed that the thermal conductance was easily tuned by a factor of three. Moreover, the highly linear performance allows for the method to be used as a technology to produce adaptive IR detectors with high dynamic range. However, despite its better tunability, in-use stiction was observed in the devices with stoppers and thus in-depth repeatability studies were not possible. To prevent in-use stiction, the vertical gap between the stoppers and the microplate can be increased, or the stopper could be fabricated with a material having low adhesion. References [1] MCS640 Thermal Imager Operator’s Manual, Lumasense Technologies, 2011. [2] W.B. Song, J.J. Talghader, Design and characterization of adaptive microbolometers, J. Micromech. Microeng. 16 (2006) 1073–1079. [3] V.N. Leonov, D.P. Butler, Two-color thermal detector with thermal chopping for infrared focal-plane arrays, Appl. Optics 40 (2001) 2601–2610. [4] W.B. Song, J.J. Talghader, Adjustable responsivity for thermal infrared detectors, Appl. Phys. Lett. 81 (2002) 550–552.
[5] N. Topaloglu, P.M. Nieva, M. Yavuz, J.P. Huissoon, A pixel-by-pixel thermal conductance tuning mechanism for uncooled microbolometers, in: Microsystems and Nanoelectronics Research Conference, 2008. MNRC 2008, 1st, 2008, pp. 201–204. [6] F. Niklaus, C. Vieider, H. Jakobsen, Mems-based uncooled infrared bolometer arrays: a review, in: Proc. SPIE, vol. 6836, 2007, pp. 68360D–68360D-15. [7] J. Carter, A. Cowen, B. Hardy, R. Mahadevan, M. Stonefield, S. Wilcenski, POLYMUMPs Design Handbook Revision v11.0, 2005. [8] R.N. Supino, J.J. Talghader, Average optical power monitoring in micromirrors, IEEE J. Sel. Top. Quant. Electron. 8 (2002) 12–18. [9] W. Rohsenow, J. Hartnett, Y. Cho, Handbook of Heat Transfer, 3rd ed., McGrawHill, New York, 1998. [10] M.R. Sridhar, M.M. Yovanovich, Review of elastic and plastic contact conductance models: comparison with experiment, J. Thermophys. Heat Trans. 8 (1994) 633–640. [11] F.H. Milanez, M.M. Yovanovich, J.R. Culham, Effect of surface asperity truncation on thermal contact conductance, IEEE Trans. Compon. Pack. Technol. 26 (2003) 48–54. [12] F.H. Milanez, M.M. Yovanovich, M.B.H. Mantelli, Thermal contact conductance at low contact pressures, J. Thermophys. Heat Trans. 18 (2004) 37–44. [13] L.M. Phinney, G. Lin, J. Wellman, A. Garcia, Surface roughness measurements of micromachined polycrystalline silicon films, J. Micromech. Microeng. 14 (2004) 927–931. [14] P. Eriksson, J.Y. Andersson, G. Stemme, Thermal characterization of surfacemicromachined silicon nitride membranes for thermal infrared detectors, J. Microelectromech. Syst. 6 (1997) 55–61. [15] D.P. Neikirk, W.W. Lam, D.B. Rutledge, Far-infrared microbolometer detectors, Int. J. Infrared Millim. Waves 5 (1984) 245–278. 10.1007/BF01009656. [16] C.M. Hanson, Uncooled IR detector performance limits and barriers, in: Proc. SPIE, vol. 4028, Orlando, FL, 2000, pp. 2–11.
Contents lists available at SciVerse ScienceDirect
Infrared Physics & Technology journal homepage: www.elsevier.com/locate/infrared
An effective thermal conductance tuning mechanism for uncooled microbolometers Nezih Topaloglu a,⇑, Patricia M. Nieva b, Mustafa Yavuz b, Jan P. Huissoon b a b
Yeditepe University, Department of Mechanical Engineering, 34755 Istanbul, Turkey University of Waterloo, Department of Mechanical and Mechatronics Engineering, 200 University Avenue West, Waterloo, ON, Canada N2L 3G1
h i g h l i g h t s " A thermal conductance tuning mechanism for uncooled microbolometers is proposed. " A pixel-by-pixel tuning is realized and a stopper mechanism is used. " Thermal conductance can be tuned by a factor of three. " The tunability is highly linear.
a r t i c l e
i n f o
Article history: Received 28 May 2012 Available online 3 January 2013 Keywords: MEMS Uncooled microbolometers Thermal modeling Tunable thermal conductance
a b s t r a c t With the increasing demand on infrared (IR) detectors for imaging harsh environment processes, widening the application range of uncooled microbolometer arrays has become an important research area. An efficient way of increasing this range is tuning the thermal conductance of the microbolometer array using electrostatic actuation, which is usually achieved by directly applying an actuation voltage to the substrate. However, this method does not allow pixel-by-pixel actuation, limiting the tunability. In this paper, we present a new method of actuation which uses the micromirror located below the microbolometer as the actuation terminal. We demonstrate that using micromirror actuation, the thermal conductance can be tuned by a factor of three. An analytical model to calculate the thermal conductance of this new type of microbolometer is presented. Results of the model are compared to finite element simulations and experimental measurements on a test structure fabricated for this purpose, showing good agreement. The new tuning mechanism provides a fairly linear thermal conductance tunability, thus making it a promising thermal conductance controlling mechanism for adaptive IR detectors. Ó 2012 Elsevier B.V. All rights reserved.
1. Introduction The demand for uncooled microbolometer focal planar arrays (FPAs) has increased significantly over the last few years, due to their high sensitivity and low cost. Due to the recent growth in the application areas of thermal imagers, such as high temperature process imaging [1], increasing the device dynamic range has become one of the main challenges in microbolometer research. For instance, when a high temperature process (e.g. 3000 K during an exothermic reaction) is imaged with a microbolometer FPA, the high radiation power can increase the temperature of the device by about 100 °C [2]. This increase can lead to a number of undesirable effects, such as nonlinear response, plastic deformation of both the arms and the microplate, and ultimately device failure. To overcome these problems, the thermal conductance ⇑ Corresponding author. E-mail addresses: [email protected] (N. Topaloglu), pnieva@ uwaterloo.ca (P.M. Nieva). URL: http://simslab.uwaterloo.ca (P.M. Nieva). 1350-4495/$ - see front matter Ó 2012 Elsevier B.V. All rights reserved. http://dx.doi.org/10.1016/j.infrared.2012.12.039
(G) can be varied directly by contacting the detector with its substrate using electrostatic actuation [3,4], where the thermal conductance of the microplate is controlled by the thermal contact conductance (TCC) of the detector-substrate interface [4]. However, thermal conductance is a parameter that directly affects the responsivity and the response time of a microbolometer. As the thermal conductance decreases, both the responsivity and the response time increase. Hence, any variation in thermal conductance results in a trade-off between the responsivity and response speed. Therefore, a microbolometer FPA with adjustable thermal conductance can give the user the opportunity to switch from a high temperature resolution and low frame rate mode to a low resolution and high frame rate mode, depending on the application. Previous thermal conductance tuning mechanisms apply a certain voltage directly to the substrate to actuate the microplate [3,4]. Nonetheless, this type of actuation does not allow pixel-bypixel thermal conductance tuning. In addition, due to the large contact area at pull-in, the change in thermal conductance is sudden and very high, which prevents linear thermal conductance tunability.
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Previous research work in our group presented a new microbolometer design that used a stopper mechanism to adjust the thermal conductance [5]. The stopper mechanism allowed for a fairly linear pixel-by-pixel thermal conductance tunability due to the small contact area at pull-in. In this same work, it was demonstrated that the thermal conductance for a 120 120 lm2 microplate size microbolometer was doubled, with respect to the unactuated state, when an actuation voltage of 12 V was used. In this paper, we propose an improved microbolometer design which also includes a stopper mechanism. The microplate size of the improved microbolometer is 80 80 lm2 which is considerably smaller when compared to the 120 120 lm2. This size is also closer to the sizes of commercial microbolometers which range from 17 17 lm2 to around 40 40 lm2 [6]. In addition, the arm length is increased from 160 lm to 180 lm, for increasing the thermal isolation and to decrease the contact voltage. The design and fabrication of the proposed tunable thermal conductance mechanism are presented in Section 2, and then compared to existing mechanisms. Section 3 includes an analytical thermal conductance model, which calculates the thermal conductance of the device in terms of the actuation voltage. A comparison of the analytical results, with finite element simulations, electrical capacitance measurements and thermal conductance measurements is presented in Section 4. Finally, the conclusions are drawn in Section 5. 2. Design and fabrication Fig. 1 shows a schematic of a tunable thermal conductance mechanism for microbolometers, demonstrated by Song and Talghader in 2002 [4]. In this design, the microplate is actuated by applying a bias voltage to the substrate (Fig. 1b). In this case, the authors demonstrated that thermal conductance can be increased up to four times when the detector snaps down completely onto the substrate. Later, the same authors demonstrated that the tunability range of the thermal conductance can be increased with a similar improved design, which used the same substrate biasing method [2]. Although biasing the substrate adds great tunability to the thermal conductance, it has a disadvantage. The bias voltage applied to the substrate results in the actuation of the whole microbolometer pixel array, which makes pixel-by-pixel thermal conductance tuning difficult. This problem could be avoided by biasing the
Fig. 1. Cross-sectional schematic of the microbolometer thermal conductance tuning mechanism demonstrated in [4] in (a) unactuated state and (b) actuated state. (Not drawn to scale.)
micromirror plate, which is located below the IR sensitive microplate, instead of the substrate. Since the micromirror is used only for increasing the infrared absorption of the microbolometer pixel; the effect on the normal operation of the microbolometer when a voltage is applied to it will most likely be negligible. However, to maintain a continuous electrostatic actuation force and safe operation, there should not be any electrical contact between the microplate and the micromirror. This can be achieved by preventing the snap-down of the microplate onto the micromirror once a voltage higher than the pull-in voltage is applied. Depositing a sufficiently thick non-conductive material on the microplate that can electrically isolate the micromirror could be a potential solution. However, in this case, the thermal conductance will reach its maximum value abruptly, resulting in a digitized one-step thermal conductance tuning. To precisely control the thermal conductance according to the environmental conditions, a thermal conductance that changes linearly with the applied voltage is required. To overcome the problems mentioned above, a stopper mechanism was designed and it is illustrated in Fig. 2. In this design, the contact between the microplate and the micromirror is avoided using stoppers. Tuning the thermal conductance is achieved using electrostatic actuation of the stoppers. The stoppers are rigid structures anchored to the substrate, and they extend into the gap between the microplate and the base plate. During operation, as the microplate deflects sufficiently towards the biased micromirror, the stoppers prevent the upper plate from deflecting further upon contact, thus avoiding the contact with the micromirror (see Fig. 2b). At the same time, the thermal conductance increases, since the heat flows to the substrate mainly through the stoppers, which provide a good heat link between the microbolometer and the substrate. When the bias voltage is decreased to 0 V, the microplate reverts back to its original position, thus eliminating the heat link provided by the stoppers (Fig. 2a). In that case, thermal conductance will decrease back to its lowest value, which is then only limited by heat conduction through the microbolometer arms. To demonstrate this stopper mechanism, several test structures were fabricated using PolyMUMPs (Polysilicon Multi-User MEMS Process) [7]. PolyMUMPs is a three-layer polysilicon process, which has well-defined process steps and design rules. Fig. 3 shows the SEM image of a fabricated test structure. Since the stopper should extend in between the microplate and the base plate (mirror), the base plate, the stoppers and the microbolometer were made by the Poly 0, Poly 1 and Poly 2 layers, respectively. The air gap between the microplate and micromirror is 2.75 lm. It should be noted that in an actual microbolometer, the micromirror is made of metal.
Fig. 2. Cross-sectional schematic of the microbolometer with the proposed stopper, in (a) unactuated state and (b) actuated state. (Not drawn to scale.)
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At the unactuated state, the stoppers do not have any effect on the thermal conductance, and the conduction through the microbolometer arms is the dominant heat conduction mechanism. Therefore, the thermal conductance can be calculated by applying Fourier’s law of conduction to the arms.
Gunact ¼
2kPolySi wPolySi t PolySi 2kgold wgold tgold þ ; LPolySi Lgold
ð1Þ
where k, w, t and L are the thermal conductivity, width, thickness, and length of the corresponding arm layer, respectively. At the actuated state, the microplate contacts the stoppers, forming a new heat link between the microplate and the substrate. The temperature decrease at this heat link is mainly determined by the thermal contact conductance between the microplate and the stoppers. Therefore, the thermal conductance at the actuated state (Gact) can be stated as
Gact ¼ Gunact þ Gc ;
Fig. 3. Scanning electron microscope (SEM) image of the test structure used in this paper: a 80 lm 80 lm microplate, connected to the substrate with two 180 lm 10 lm arms.
However, for these test structures, the Poly 0 layer is used as the base plate of the electrostatic actuator instead of a metallic layer, due to the limitations of the PolyMUMPs process. Nevertheless, this situation is not expected to affect the experimental results, since the micromirror material has no effect on the electrostatic actuation. The authors would also like to acknowledge here that the addition of stoppers results in added complexity and a higher pixel size, which might not be an option for some FPA applications. However, a compromised solution could result from further optimization of the shape, materials and dimensions of the microbolometer pixel, specially for high temperature applications such as the ones mentioned before. This however is not within the scope of this paper. The area of the microplate of the test structure used for the purpose of this paper is 80 80 lm2. This area is considerably smaller than the microplate area of the test structure presented in Ref. [5], which showed a similar performance. It should be noted that stateof-art microbolometer designs can have microplate areas as small as 17 17 lm2 [6], thus reducing the size of a microplate while maintaining the performance of new microbolometer designs could be used as a measure of improvement. To improve and prevent stiction during final release, 7 lm 7 lm holes and 4 lm 4 lm dimples are used. The microbolometer arms are 180 lm long and 10 lm wide and were coated with the Au metallization, to make sure that the resistive heating occurs mainly within the microplate. 3. Thermal conductance modeling In this section, modeling of the thermal conductance of the proposed mechanism is presented. The main heat transfer mechanisms involved in the operation of a microbolometer are convection and radiation between the pixel and its surroundings, and conduction to the substrate. For thermal isolation, microbolometers are packaged in vacuum, which means convection can be neglected. Radiation can also be assumed to be negligible, since the pixel never reaches extremely high temperatures during operation [8]. Therefore, for both the unactuated and the actuated state, the thermal conductance models, for our particular case, are based on the conduction of the heat to the substrate through the arms and the stoppers.
ð2Þ
where Gc is the thermal contact conductance between the microplate and the stoppers. The term Gc stems from the fact that the contacting surfaces are not perfectly smooth and flat. Therefore, as the two surfaces contact, they will touch only at a portion of the apparent contact area, forming microgaps between them. As a result, the heat is transferred only through the contacting asperities, which results in a temperature drop across the contact. A small portion of a contact between two nominally flat surfaces (also called conforming rough surfaces) is shown schematically in Fig. 4a. Each surface has a mean plane, which is shown in the figure with a dotted-dashed line. The surface roughness and the absolute mean asperity slope of the surface i are defined by ri and mi, respectively, both of which can be determined from the corresponding surface profile data as [9]
ri
sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Z 1 L 2 ¼ y ðxÞdx; L 0 i
mi ¼
1 L
Z 0
L
dyi dx; dx
ð3Þ
ð4Þ
where yi(x) is the distance of the points on surface i from the mean plane, and L is the length of a trace. The contact between the rough surfaces can be transformed into the contact between a flat surface and a rough surface, where the effective roughness, r, and the effective absolute mean asperity slope, m, are defined as [9]
r¼
qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r21 þ r22 ;
ð5Þ
m¼
qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi m21 þ m22 :
ð6Þ
A small portion of the contact after this transformation is shown in Fig. 4b. Gc depends on many parameters, such as the surface roughness and the contact pressure. Various models are available for determining Gc [9]. Most of these models are based on the assumption that surface asperities are distributed randomly over the contact area, and they have Gaussian height distribution about some mean plane passing through each of them. These models are called fully Gaussian (FG) models. However, it was reported that at low contact pressures, the fully Gaussian models systematically underestimate the thermal contact conductance [10]. An explanation to this inaccuracy at low contact pressures is proposed by Milanez et al. [11]. There, it was reported that the real surfaces may have Gaussian surface distribution up to around 4.5r, however they generally do not have asperities whose height is above 4.5r [11]. Therefore,
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Fig. 4. Schematic showing a small portion of (a) a contact formed between two conforming rough surfaces, (b) the same contact transformed into the contact of a flat and a rough surface. Reprinted from [9].
the distance between the mean planes at low contact pressures is shorter than expected and the contact conductance is greater than expected. As the contact pressure increases, more number of shorter asperities come into contact, which explains the accuracy of the fully Gaussian model at higher pressures. As a solution to this inaccuracy, a truncated gaussian (TG) model was developed [11,12]. This model assumes that the heights of the surface asperities follow the Gaussian distribution up to a defined value of ztrr, where ztr is the relative truncation level. Due to its higher accuracy in low contact pressures, the TG model is employed for determining Gc. Assuming the asperities deform elastically upon contact, Gc can be expressed as [12]
sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi m 1 ; Gc ¼ 1:25Aa ks ðP=He Þ0:95 ½1 þ 1=f 0:9289 1 1þf r
ð7Þ
where
pffiffiffiffiffiffiffi f ¼ ðP=He Þ 2p ztr exp z2tr =2 :
ð8Þ
In Eqs. (7) and (8), Aa is the apparent contact area, P is the contact pressure and He is the effective elastic microhardness, given by
m E ; He ¼ pffiffiffi 2 2 2 1m
ð9Þ
where E and m are the Young’s modulus and the Poisson’s ratio of polysilicon, respectively. To find P, the net force Fnet acting on the microplate upon contact should be determined. Approximating the microbolometer to a moving plate capacitor, where the moving plate is the microplate, Fnet is given by
F net ¼ keq ðg 0 gÞ
0 AV 2act 2g 2
area, P can be found. The contact voltage Vcontact, which will be defined here as the voltage where the microplate touches the stoppers, also needs to be found, to determine the beginning of the actuated state. Since the height of the stopper is smaller than the height at which pull-in occurs, Vcontact can be found by equating Fnet to zero, and g to 2g0/3, which yields
V contact
sffiffiffiffiffiffiffiffiffiffiffiffiffiffi 8keq g 30 ¼ : 270 A
ð11Þ
4. Experiments and discussion 4.1. Optical profiler measurements A Wyco NT100 optical profiler was used to determine the thicknesses of the air gap and the polysilicon films forming the test pixel shown in Fig. 3 after fabrication. They were also used to determine the effective surface roughness parameters (r and m), and the surface truncation parameters (ztr). The parameters, extracted from the optical profiler data are tabulated in Table 1. The optical profiler data shows that the gap between the microplate and the base plate is 2.7 lm at the unactuated state and 1.7 lm at the actuated state. The gap thickness of 1.7 lm is less than the design value of the Oxide 1 layer (2 lm [7]), which shows
Table 1 The thicknesses and roughness parameters extracted for the test structure shown in Fig. 3 using data from an optical profiler. Thicknesses (lm)
;
ð10Þ
where A is the microplate area, g0 is the initial gap thickness, and g is the final gap thickness, 0 is the permittivity of free space, and keq is the equivalent spring constant of the arms, which can be found by beam bending analysis. Dividing Fnet to Ac, which is the total contact
Poly 0 Poly 1 Poly 2 Gold Initial gap (g0) Gap after contact (g)
Roughness parameters 0.6 2.0 1.3 0.5 2.7 1.7
r1 m1
r2 m2 ztr
12.14 nm 0.042 2.06 nm [13] 0.007 2.44
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that the microplate successfully contacts the stoppers at the actuated state. In Figs. 5 and 6, the 3D surface profile and X and Y profile of the test structure captured by WYKO NT100 Optical Profiler is plotted, when the actuation voltage is 12 V. In a parallel plate capacitor, pull-in occurs when the gap between the capacitor plates is less than two-thirds of the initial gap. In our case, the results of the optical profiler measurements show that the gap after pull-in (g) is less than two-thirds of the initial gap (g0) value. Hence, Eq. (11) can be used as an analytical expression to calculate the contact voltage. To calculate r and m, the surface roughness of the contacting surfaces should be determined. Taking the top surface of the stoppers (Poly 1) as surface 1, r1 and m1 is calculated from the optical profiler data as r1 = 12.14 nm and m1 = 0.042. The surface roughness parameters of the other contacting surface (r2 and m2), which is the bottom surface of the microplate (Poly 2), cannot be measured by the optical profiler. Instead, r2 is taken from [13] as 2.06 nm and m2 is estimated according to the formula m2 = m1(r2/r1). Using these values, r and m can be found (r = 12.31 nm, m = 0.043). The truncation level ztr is determined by extracting the height of the highest asperity from the surface profile data and dividing it to the surface roughness, which yields ztr = 2.44. 4.2. Finite element model simulations To determine the contact voltages of the test structures in the design stage, a finite element model (FEM) was constructed using the commercial finite element analysis (FEA) software ANSYSTM. The total number of elements of the constructed model was around 100000. The structure was modeled with the 3-D 20 node structural SOLID 95 element, whereas the air surrounding the structure was modeled with the electrostatic SOLID 122 element. The mechanical properties of polysilicon and gold were taken from the PolyMUMPs manual [7]. For simplicity, the stoppers were not included in the FEM. After the construction of the model, a coupled electrostatic-structural solution was performed using the ESSOLV macro. The electrostatic solver calculates the mechanical forces based on the potential difference, and the structural solver uses the force distribution to calculate the deflection. If convergence is not reached after the first solution step, the deflection data is used again by the electrostatic solver to recalculate the force distribution. This loop continues until convergence is reached. Using the gap and thin-film thicknesses obtained from the optical profiler data (Table 1), the contact voltage calculated using the FEM model was 13.5 V. The theoretical contact voltage, i.e. the contact voltage calculated by Eq. (11), is found as 12.4 V, which is fairly close to the contact voltage found by the FEM. The small discrepancy between these values are mainly attributed to the limitations introduced in the theoretical model when assuming a perfectly flat microplate and when calculating the equivalent spring constant of the arms using Euler Bernoulli beam theory.
Fig. 5. 3D image of the test structure created by WYKO NT100 Optical Profiler.
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4.3. Electrical capacitance measurements A reliable way to measure the contact voltage is to measure the electrical capacitance. As the actuation voltage increases, the electrical capacitance increases as the result of the decrease in gap (g). A sudden increase is expected once contact voltage is reached. The voltage versus capacitance plot for the device under study, measured with an Agilent E4980A Precision LCR Meter is shown in Fig. 7. As can be seen from Fig. 7, a jump in the capacitance is observed roughly at 11 V, which is the voltage at which the microplate touches the stoppers. The increase in capacitance after 12 V is less than 0.2%, which for the effects of this paper can be assumed as negligible. The authors believe that compared to the FEM (13.5 V) and theoretically calculated contact voltage (12.4 V), the lower experimental contact voltage is mostly due to the effect of the adhesion force between the stoppers and the microplate, which was not considered in the FEM and theoretical contact voltage calculation. 4.4. Thermal conductance measurements The thermal conductance was measured under vacuum (<4 mTorr) using the Joule heating method [14]. In this method, heat is generated by biasing the microplate with a voltage or current source under vacuum. By measuring the current at each voltage step, a plot of dissipated power versus resistance is obtained. G is calculated from the slope of this curve, according to the following formula [15]
R ¼ R0 þ
1 R0 aI2 R; G
ð12Þ
where I is the current, R0 is the resistance at 300 K and a is the temperature coefficient of resistance (TCR), found as 5.2 104 K1 by obtaining the temperature versus resistance plot in a temperature controlled environment. The power versus resistance measured for the microbolometer pixel shown in Fig. 3, is plotted in Fig. 8. In the unactuated state, G is found to be 1.29 105 W/K. This value is in good agreement with 1.30 105 W/K, which is the thermal conductance calculated by Eq. (1). With an actuation voltage of 11 V, the thermal conductance increases to 2.44 105 W/K. As discussed previously, this increase in G is expected, due to the new heat links formed by the contact of microplate to the stoppers. As the actuation voltage is further increased, the pressure of the microplate on the stoppers increases, resulting in higher TCC. This explains the increase in thermal conductance to 3.71 105 W/K, with an actuation voltage of 35 V. After around 35 V, the devices failed mostly due to overheating, as a result of breakdown failure. It should be noted that due to the limitations of the PolyMUMPs process, the dimensions of the fabricated structures are not ideal (e.g. arms cannot be as thin and materials cannot be changed), so thermal conductances and contact voltages are not low enough for it to be used as a commercial microbolometer [16]. By using a custom fabrication and state-of-the-art fabrication equipment, lower actuation voltages and higher tunability in thermal conductance are expected. The tunability range of the thermal conductance can be found by dividing the thermal conductance at its highest value as a result of applied voltage, to thermal conductance at the unactuated state. In Fig. 9, the measured thermal conductance is compared to calculated values using Eq. (2) of the thermal conductance model presented in the previous Section. From Fig. 9, it can be concluded that the results obtained using the analytical model shows a similar trend when compared to the ones obtained experimentally. However, it underestimates the thermal conductance at the
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Fig. 6. Profile of the test structure in (a) X and (b) Y directions.
Fig. 7. C–V response of the test structure shown in Fig. 3 measured by an LCR meter. Note that there is a sudden jump in the capacitance roughly at 11 V.
Fig. 8. A plot of the power versus resistance, for the test structure with stoppers shown in Fig. 3. The top curve corresponds to the unactuated state (Vact = 0).
actuated state. In addition, due to the discrepancy in the experimental and theoretical contact voltage, the onset of the actuated state of two curves differ by around 2 V. The inaccuracy in the contact voltage estimation also leads to an inaccuracy in contact pressure estimation, which we can be attributed as the main reason of the underestimation of thermal conductance.
To observe the effect of stoppers on the contact voltage and the thermal conductance, Fig. 10 shows the thermal conductance (G) versus actuation voltage of two different microbolometer structures. One of these structures is the microbolometer in Fig. 3 (structure A). The second structure is a microbolometer that has the same geometry and has been fabricated in the same batch
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Fig. 9. Thermal conductance versus actuation voltage of the test structure. The solid curve corresponds to the experimental data and the dashed curve corresponds to the analytical thermal contact model.
Fig. 10. Thermal conductance versus actuation voltage of test structures with and without the stopper layer.
using PolyMUMPs. This device still keeps the polysilicon micromirror at the bottom but does not have the stoppers (structure B). To avoid the electrical contact between the actuation terminals, structure B is actuated by biasing the substrate, which is isolated from the surface structures by a 600 nm silicon nitride layer. As it can be observed, before contact, both structures exhibit a constant thermal conductance (1.29 105 W/K for structure A and 1.25 105 W/K for structure B). At contact voltage, the thermal conductance of structure B jumps from 1.25 105 W/K to 9.51 105 W/K, which corresponds to a tunability range of around 8. On the other hand, the thermal conductance of structure A at contact becomes 2.44 105 W/K. This is expected since at contact, the microplate in structure B snaps down to the substrate, resulting in a higher contact area than that on structure A at contact. In addition, when the structure B is at contact, the gap between the microplate and the substrate is smaller (1.1 lm corresponding to the nitride and micromirror layers, refer to Fig. 2), compared to the contact of structure A, resulting in a higher electrostatic attraction. From Fig. 10, we can also see that for the test structure B, as the actuation voltage is increased from 0 to the contact voltage, there is less than a 0.3% change in G. Therefore, the thermal conductance can only be tuned to two different values, the one before, and the one after contact. In comparison, the G of the structure with stoppers increases linearly at a rate of 5 105 W/K/V, due to the increase in TCC between the stoppers and the microplate. This linear response provides us with a continuous tunability range as well as the ability to control the thermal conductance more precisely. The tunability range of test structure B is calculated to be around 3. From Fig. 10, it should be also inferred that the contact voltage of structure B is measured to be higher than that of structure A, which shows the effect of stoppers-microplate adhesion force on the contact voltage. As it was mentioned before, a common problem that may arise during electrostatic actuation is the in-use stiction. Although the
devices type A are expected to exhibit low adhesion forces due to their smaller contact area, we observed in-use stiction during testing, when the voltage reaches contact voltage. On the other hand, no evidence of in-use stiction was observed in the structures type B. This may be attributed to the fact that the stoppers limited the deflection of the microplate, which resulted in a restoring force smaller than the adhesion force, when the actuation voltage was removed. We believe that in order to avoid stiction, the gap between the stoppers and the microplate should be higher. This way, the microplate will need to deflect more before contacting the stoppers, which in turn will increase the restoring force. An adequate increase in the restoring force will easily bring the microplate back to its original position when the actuation voltage is removed. This can be achieved if the stoppers are fabricated closer to the base plate by using a customized fabrication process instead of PolyMUMPs. In addition, since the stopper material affects the adhesion force, choosing an appropriate stopper material could further decrease this force. The problems encountered with inuse stiction during experimental testing prevented us from conducting repeatability tests using the same device. However, onset actuation voltages measured in identical test structures as the one shown in Fig. 3, from the same batch but different chips, were also around 11 V.
5. Conclusion An efficient thermal conductance tuning mechanism for uncooled microbolometer, based on electrostatic actuation has been proposed. Instead of biasing the substrate and to realize a pixelby-pixel tuning, a stopper mechanism is used to prevent electrical contact between the electrostatic actuation terminals. We demonstrate that this mechanism linearizes the thermal conductance tuning after the onset actuation voltage needed to contact the microplate with the stoppers. A thermal contact conductance
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model was used for estimating the thermal conductance as a unction of the actuation voltage. Using this stopper mechanism, the thermal conductance of the microbolometer was increased with a linear tunability ranging from 2.44 105 W/K to 3.71 105 W/K. Considering the lowest and highest thermal conductance measured (1.29 105 W/K and 3.71 105 W/K), we observed that the thermal conductance was easily tuned by a factor of three. Moreover, the highly linear performance allows for the method to be used as a technology to produce adaptive IR detectors with high dynamic range. However, despite its better tunability, in-use stiction was observed in the devices with stoppers and thus in-depth repeatability studies were not possible. To prevent in-use stiction, the vertical gap between the stoppers and the microplate can be increased, or the stopper could be fabricated with a material having low adhesion. References [1] MCS640 Thermal Imager Operator’s Manual, Lumasense Technologies, 2011. [2] W.B. Song, J.J. Talghader, Design and characterization of adaptive microbolometers, J. Micromech. Microeng. 16 (2006) 1073–1079. [3] V.N. Leonov, D.P. Butler, Two-color thermal detector with thermal chopping for infrared focal-plane arrays, Appl. Optics 40 (2001) 2601–2610. [4] W.B. Song, J.J. Talghader, Adjustable responsivity for thermal infrared detectors, Appl. Phys. Lett. 81 (2002) 550–552.
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