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Influence of temperature on the actuation voltage of RF-MEMS switches
Microelectronics Reliability xxx (2013) xxx–xxx
Contents lists available at SciVerse ScienceDirect
Microelectronics Reliability journal homepage: www.elsevier.com/locate/microrel
Influence of temperature on the actuation voltage of RF-MEMS switches V. Mulloni ⇑, F. Solazzi, F. Ficorella, A. Collini, B. Margesin Fondazione B. Kessler, Center for Materials and Microsystems, Via Sommarive 18, I-38123 Trento, Italy
a r t i c l e
i n f o
Article history: Received 7 September 2012 Received in revised form 17 January 2013 Accepted 17 January 2013 Available online xxxx
a b s t r a c t Most of the actual applications for RF-MEMS switch require high reliability, but consolidated qualification procedures are still lacking. This paper focuses, in particular, on the role of temperature on the switch reliability from a mechanical point of view, showing how this depends on the switch architecture and membrane material. Double clamped switches are sensitive to buckling, and this is the factor limiting their operational temperature, even though the range exploitable can be wide enough for many applications. Residual stress and thermal expansion coefficient of the mobile membrane are the most important parameters to understand and control this phenomenon. Cantilever switches are less influenced by the temperature in their performances, and have a much wider operational range. Other temperature-related factors are affecting the switch reliability in this case, such as elastic modulus variation, dielectric charging effects and creep. Ó 2013 Elsevier Ltd. All rights reserved.
1. Introduction In the last years RF-MEMS switches have been developed by several research groups and companies, demonstrating outstanding performances in terms of RF-characteristics in a broad range of frequencies. Nonetheless RF-MEMS switch reliability is still an open issue which so far prevented an extensive use of these devices. In applications where temperature variations are important, such us in satellite communications, or in the packaging process itself, the effect of temperature on device performances is of great importance for evaluating device reliability, and usually a defined operational temperature interval is explicitly required to qualify a defined switch technology. Temperature can affect switch characteristics in several ways. Electrical polarization mechanisms due to charging, such us free charge redistribution, charge injection and charge dissipation are temperature dependent [1,2], but the mechanical properties of the switching membrane are also strongly affected by the temperature [3], and the combined effects of all these physical variations can couple and superimpose increasing the complexity of reliability testing and analysis. In this view, the data presented in this paper focuses only on the mechanical modifications induced by the temperature on the MEMS mobile structure, with the purpose to separate these effects from other types of temperature induced reliability issues when analyzing switch reliability data and trying to design accelerating stress tests. Usually, after the release process, the mobile part of MEMS switch is deformed to some extent in the out-of-plane direction, due to a complicated interplay of its residual stress and stress ⇑ Corresponding author. E-mail address: [email protected] (V. Mulloni).
gradient and the modifications induced by the release process itself [4]. This out-of plane profile may sensibly change with temperature, and needless to say, the actuation properties of the device are strongly influenced by these variations. The key mechanical parameter to understand these modifications is however not the same in all type of switches, and in particular, is quite different in the case of double-clamped or single clamped switches, which are the most common switch typologies. In this paper the temperature-induced actuation voltage variations in the temperature interval 25–100 °C are measured, interpreted and discussed for a double-clamped and a cantilever switch. Even though the switching membrane are made of the same material, the operational range and the actuation voltage variations are very different for the two typologies. The main reasons for this differences are described, resulting in a stress dominated behavior in the case of double-clamped switch, while other factors are affecting the behavior of cantilever switch. 2. Temperature related mechanical variations 2.1. Double-clamped structures In the case of a switch membrane clamped at both ends, the key mechanical parameter influenced by the temperature is the residual stress of the membrane. The usual analytical formula reported for the actuation voltage Vpi of a switch is [5]:
sffiffiffiffiffiffiffiffiffiffiffiffiffi 3 8kd0 V pi ¼ 27e0 A
ð1Þ
where d0 is the zero-voltage gap spacing, A the electrode area, e0 is the vacuum dielectric permittivity and k is the spring constant.
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Please cite this article in press as: Mulloni V et al. Influence of temperature on the actuation voltage of RF-MEMS switches. Microelectron Reliab (2013), http://dx.doi.org/10.1016/j.microrel.2013.01.007
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While d0, A, and e0 can be considered temperature independent within a wide range, the spring constant is strongly temperature dependent [6]. In a double-clamped structure, k is determined by two main contributions [7]: one is given by the elastic response to an applied force and is proportional to the Young’s modulus of the structure, while the second part is determined by the residual stress of the membrane. In the most common case, the switch substrate is silicon or quartz, and the mobile membrane is made of metal. In this case, being the membrane anchored to the substrate, the different thermal expansion coefficients produce an additional tension in the membrane when the temperature is lowered with respect to the ambient temperature, and a compression when the temperature is raised. The membrane itself usually possesses an intrinsic stress that may strongly vary with processing conditions, but it is normally tensile. The switch actuation voltage decreases with decreasing the tensile membrane stress, that is it decreases with increasing temperature until it reaches a minimum value, while its shape does not change and the gap remains constant. After this minimum value, a discontinuity appears because the membrane buckles and the switch can no longer properly actuate. This behavior can be represented by the following equation:
V 2pi / k ¼ kelastic ðEÞ þ B½rresidual
Tamb
Eðametal asubstrate ÞDT
ð2Þ
where DT is the temperature difference with respect to ambient, which is taken as reference temperature, E is the effective Young’s modulus of the metal, ametal and asubstrate are the thermal expansion coefficients of the metal and the substrate respectively, and B is a convenient geometric factor which depends on the switch geometry and dimensions. The minimum actuation voltage is reached when the residual stress of the membrane rresidual_Tamb is perfectly compensated by the last term of Eq. (2), which is the thermal stress. At the temperature in which this occurs the membrane is stress free, and its actuation voltage is determined only by its geometry and by the elastic modulus of the metal. However, the stress interval between zero stress and buckling is very narrow [6], as can be estimated by the Eulero’s formula. For practical purposes it is correct to assume that as soon as the stress turns compressive, the membrane buckles. For double-clamped switches then, the range of operational temperatures is then defined by the maximum acceptable actuation voltage (lowest temperature) and by the buckling temperature (highest temperature). Other factors can play a role, such as the variation of Young’s modulus with temperature, which decreases with increasing temperature, adding its effect to that of thermal stress, and charging phenomena, that may become faster at higher temperature [2]. However, they are estimated to be much less relevant with respect to the stress contribution. 2.2. Single-clamped structures Single clamped switches are of cantilever type and in first approximation can be considered stress free, because they can expand or contract in response to temperature variations. This means that the term within square brackets in Eq. (2) vanishes and k = kelastic(E). Ruling out the major temperature-related source of Vpi variation, other effects such as charging, Young’s modulus variation and, when the temperature is high enough, plastic relaxation and creep phenomena may become visible. The role of elastic modulus variation can be understood considering that in first approximation k depends linearly on E [7], and then, in force of Eq. (1), Vpi depends on the square root of E. Moreover, even though for this switch typology the average biaxial stress is virtually zero, many cantilever structures possess a relevant stress gradient which results in a marked non-planarity of the suspended membrane. This gradient develops during the release phase [4] and could be temperature dependent. In this case
the actuation voltage can be affected by the gradient variation resulting in a change in the average actuation gap. This effect, when present, can be potentially important, given the marked dependence of Vi from the average actuation gap d0 in Eq. (1). As reported before, all the mentioned effects are estimated to be much less severe than the stress variation in the case of clamped–clamped switches, and then the operational temperature range of cantilever structures is predicted to be much wider than that of clamped–clamped membranes. In particular the maximum operational temperature, when high enough, is likely to be determined by creep [8], which is strongly temperature dependent and may become sensible especially after prolonged actuation. The distinct contributions may however widely vary with switch design and materials.
3. Experimental The MEMS switches investigated were fabricated with an eightmask process at FBK. A detailed process description is discussed in [9], and can be briefly summarized as follows: high resistivity 525 lm thick and 4 in. wide silicon wafers were used as substrate. Actuation electrodes are made in lightly doped polysilicon, while the underpass is realized with a 650 nm-thick sputtered aluminum layer. The central underpass line is covered with 100 nm of PECVD oxide and then with 150 nm of a floating potential metal, which is made by e-gun evaporated gold. The mobile membrane is made by a 2 lm thick electrodeposited gold layer, suspended over an air gap of 2.7 lm, obtained by burning in oxygen plasma a sacrificial photoresist spacer. A second and thicker (about 3.5 lm) electrodeposited gold layer is present in the coplanar waveguide and in the anchoring region. In the case of the double clamped capacitive switch, the role of the floating metal electrode is to ensure a good electrical contact with the mobile membrane and the dielectric allowing a very high and predictable Con/Coff ratio [10,11], which is around 200 in the selected design, reported in Fig. 1a. In addition, the dielectric is removed from the actuation pads so to minimize charging phenomena, while mechanical stoppers, placed within but not in contact with the electrode area, prevent the contact between the downstate bridge and the actuation pads. For what concerns the cantilever switch, a picture of the selected device is reported in Fig. 1b. In this case the switch is ohmic, non mechanical stopper is present and the dielectric layer has not been removed above the electrode .This means that no particular design strategy has been taken to eliminate or reduce charging effects. Six polysilicon dimples covered by the underpass and by the 150 nm thick gold layer stabilize the gold–gold contact in the switching region. The switching membrane is made again with a 2 lm thick electrodeposited gold layer, while the 3.5 lm thick gold layer is present on part of the membrane to reinforce it, in order to compensate for the lower restoring force of cantilever-type switches. Surface topographies and vertical profiles were taken with a Zygo optical profiler, with a 10 magnification. The vertical resolution is about 1 nm, and the lateral resolution is 1 lm. The profiles at different temperature where taken heating the wafer at 120 °C and then monitoring the substrate temperature in the cooling phase during the measurement. It is then possible that the real temperature on the switch was few degrees above or below the measured one. Electrical measurements have been performed with an automated measurement systems made by an Accretech UF200A Wafer Prober, with a temperature controlled chuck, an Agilent 4284A Precision LCR Meter, an Agilent E5270B Measure Mainframe equipped with six Precision Source Monitor Units and an Agilent
Please cite this article in press as: Mulloni V et al. Influence of temperature on the actuation voltage of RF-MEMS switches. Microelectron Reliab (2013), http://dx.doi.org/10.1016/j.microrel.2013.01.007
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Fig. 1. Pictures of the RF-switches investigated: (a) Clamped–clamped switch. (b) Cantilever switch.
4. Results and discussion 4.1. Clamped–clamped switch The actuation voltages as a function of the chuck temperature are reported in Fig. 2, for seven identical switches distributed on different parts of the same wafer. For all the switches the actuation voltage decreases markedly with the temperature up to 70 °C, going from 42 ± 4 V down to 24 ± 2.5 V with a common trend that is similar to what expected, with the only exception of sample 1 which deviates from the expected trend already at 70 °C. The actuation voltages at 85 °C show an erratic behavior and none of the switches actuates at 100 °C. The dashed curves are the best fit obtained for each switch fitting linearly the square of Vpi as a function of the temperature in the range 20–70 °C, according to Eq. (2). It is worth noticing that the different fitting curves converge around
60
Sample 1 Sample 3 Sample 5 Sample 7
50
Actuation Voltage (V)
B2201A Low-Leakage Switching Matrix, which are all managed by an in-house developed software framework. In the case of the double-clamped capacitive switch a C–V measurement was performed with the use of the Precision LCR Meter. A bias voltage was applied between the actuation pads and the mobile bridge while the capacitance between the underpass and the bridge was measured with a signal at 1 MHz frequency and 30 mV of amplitude. The Vpi was detected when the capacitance value instantly increases from Coff to Con value. In the case of the cantilever resistive switch a R–V measurements was performed. A bias voltage was applied between the actuation pads and the mobile bridge, while the resistance between the underpass and the bridge was measured. The Vpi was detected when the bridge is connected to the underpass, allowing the current to flow and then instantly switching from an open circuit off state to a low series resistance on state. For each device and temperature, four consecutive sweeps have been generated starting from 0 V to 75 V, then from 75 V to 0 V, then from 0 V to 75 V and finally from 75 V to 0 V, without voltage interruption and with a voltage step of 0.5 V. The time required for the measurement of a single device at a given temperature was about 5 min. A recovery time of about 24 h at 25 °C was imposed between two measurements with different chuck temperatures. Temperature control was performed at wafer level by the Prober itself, with a fluctuation which is less than 1 °C. Before the measurement starts, the wafer was warmed-up at the selected temperature and allowed to thermalize, then all the devices on the wafer were sequentially measured.
Sample 2 Sample 4 Sample 6
40
30
20
10 20
30
40
50
60
70
80
90
Temperature (oC) Fig. 2. Actuation voltage as a function of the temperature for seven clamped– clamped switches on the same wafer.
80 °C, with the only exception of Sample 1, which has however been fitted only with the first three experimental points. While it is reasonable to assume that the membrane shape does not change up to 70 °C, it is straightforward to associate the unpredictable behavior at 85 °C with the emerging of buckling, and the convergence point around 80 °C can be reasonably interpreted as the temperature where the minimum actuation voltage can be reached. The data in Fig. 2 correlate well with the longitudinal profiles of the gold membrane reported in Fig. 3b, measured along the dashed lines of the surface topography of Fig. 3a. The profile at 25 °C shows that the membrane is not perfectly planar even at room temperature. The out of plane features are mainly due to the incomplete planarization of the sacrificial photoresist spacer which accommodates only partially the underneath vertical steps due to the presence of the actuation electrodes and underpass line. The underpass is actually higher than the electrodes and the stoppers due to the presence of the floating metal layer. The interesting feature visible in Fig. 3b is however the stability of the longitudinal profile in the range 20–65 °C and then the progressive enhancement of the central part of the membrane at the temperature of 85 °C and even more at 100 °C. The actuation voltages at 85 °C are in any case higher that those predicted by the fitting curves and this can be explained considering the increased effective gap as a consequence of the membrane
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Fig. 3. (a) Surface map and (b) longitudinal profiles taken along the dashed arrow in (a) for one of the clamped–clamped switches investigated.
Actuation Voltage (V)
70
60
Sample 1 Sample 3 Sample 5 Sample 7
50
Sample 2 Sample 4 Sample 6 Sample 8
40
20
30
40
50
60
70
80
90
100
25 bis
25 ter
Temperature (oC) Fig. 4. Actuation voltage as a function of the temperature for eight cantilever switches on the same wafer. Data labeled 25 bis was measured at 25 °C just after the measurement at 100 °C. Data labeled 25 ter was measured at 25 °C after a week.
bending in the upper direction. At 100 °C the bending is so pronounced that the switch cannot properly actuate because the central part of the membrane cannot touch the floating electrode underneath, being blocked by the stoppers placed on the lateral actuation electrodes. It should be noted that we cannot exclude, as theoretical possibility, a switch bending in the downward direction as the effect of buckling [12], and in this case a strong reduction of actuation voltage or stiction should be the consequence. Since this does not happen for any of the switches investigated, it is possible that the switch geometry or the presence of a stress gradient in the membrane layer forces the membrane to bend in the upward direction in the unactuated position. In fact, as will be shown further in this section, the cantilever switches fabricated in the same wafer show a marked upward bending, as reported in Fig. 5b, which indicates a markedly positive stress gradient, and this can explain the observed upward buckling. We know this gradient develops during the release phase [4], and we found it quite reproducible within the adopted switch fabrication technology from batch to batch. Some numerical analysis could give a better explanation of this experimental phenomenon, but it is beyond the purpose of this paper which reports mainly experimental data and could be a possible topic for future work. From Fig. 2 we can assume that 80 °C is approximately the temperature in which the zero stress is reached, and we can calculate
the membrane stress simply assuming that the part of Eq. (2) enclosed within square brackets is equal to zero. In the case of silicon substrate and gold membrane, ametal and asubstrate are equal to 2.3 106 K1 and 14.2 106 K1, respectively [13], while the Young’s modulus of gold is around 75 GPa. Being the temperature difference of 55 °C, the result for the residual stress at 25 °C is approximately 49 MPa. This value is very close to the stress values obtained from simulations on similar structures [14,15]. We can assume that highest operational temperature for this type of switch is then around 80 °C, and we can set the lower bound at least 20 °C for all the switches investigated, if we set an acceptable maximum actuation voltage at 60 V, using a prediction made with the curves in Fig. 2. The effective operational range is then sufficient for most of the possible application, even though the Vpi variations are quite large within this range. It is important to note that the operational range may strongly vary not only with the value of room temperature stress, but also with the type of metal used for the membrane. Residual stress, being mostly a result of processing conditions, is quite hard to control during manufacturing and, while a high stress value can raise the maximum operational temperature, it can seriously worsen other reliability problems. The choice of a metal with a low thermal expansion coefficient can be instead a big advantage [16]. For example, with a mobile membrane made of aluminum, but with the same stress value of those previously presented, the maximum operational temperature would be only 59 °C, being its Young’s modulus close to that of gold (70 GPa) but its ametal is, almost double (23 106 K1) [3], while switches made with molybdenum membrane can operate up to 150 °C [16], due to the very low thermal expansion coefficient of molybdenum (4.8 106 K1). 4.2. Cantilever switch The actuation voltages for positive bias as a function of the temperature are reported in Fig. 4 for eight cantilever switches with the design of Fig. 1b distributed on different parts of the same wafer. The dashed lines in Fig. 4 are only a guide for the eyes. The Vpi’s decrease with the temperature, but the decrease from 25 to 100 °C is limited to about 11% in all cases, and all the switches actuate up to 100 °C. The switch topography together with the profiles taken at different temperatures for one of the samples investigated are reported in Fig. 5 where two things appear clearly: first, the switch membrane bends upwards, due to an evident positive stress gradient, with a height difference from anchor to tip of 3.5 lm; second, the temperature has no visible effect on this bending within the range investigated. This rules out the gradient variation phenom-
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tip
20 oC 85 oC 100 oC
6
anchor
Height (µm)
8
4
(a)
(b)
2 0.05
0.10
0.15
0.20
0.25
Length (mm) Fig. 5. Surface map and (b) longitudinal profiles taken along the blue dashed arrow in (a) for one of the cantilever switches investigated.
ena to explain the actuation voltage decrease. A partial role in this decrease can be attributed to variations of gold elastic modulus. From the data in Refs. [17,18] this variation can be estimated of the order of 4–5% in the temperature interval 25–100 °C and then the Vpi reduction should be no more than 3%. This implies that there is some other source of temperature-induced actuation voltage change. Charging is the most likely candidate, since the switch has not been designed to minimize it, and because charging phenomena are known to accelerate with temperature [2]. It is then possible that the switch is not fully discharged when measured again at a different temperature. An evidence of the presence of charging can be given by the comparison of positive and negative actuation voltages obtained in the bipolar scan. In fact, some differences have been detected, in the range of 0.5–2 V, indicating that some charging is present, but also that the values are close to the experimental error, being the sampling interval of 0.5 V, and no clear temperature trend was detected. Furthermore, this test evidences the presence of charging but does not gives us the absolute value of this shift. Another possibility to verify the presence of charging is to measure again the Vpi at 25 °C after the measurement at 100 °C. The data obtained in this way are shown in Fig. 4 after the data at 100 °C corresponding to the label 25 bis. As can be seen, the recovery towards the initial position is minimal. This means that the measured variations are either not due only to temperature (charging or other possible but unknown effects), or are not fully reversible (creep). Creep cannot be completely excluded from data in Fig. 5b, because it can be induced by the stress of actuation, which was not been applied when the data in Fig. 5b was measured. To separate this two effects a third measurement at 25 °C was performed after a week, which is a time surely sufficient to dissipate any residual charging. The corresponding data is labeled 25 ter in Fig. 4. As can be seen, the recovery is almost complete, indicating that the measured variations are likely due to charging and no irreversible effect took place.
5. Conclusion The variation of actuation voltage with temperature was measured and discussed for two switch typologies. For clamped– clamped switches, the variations are more than 40% in the temperature interval 25–70 °C. In this case the membrane stress variation is the dominant effect, and buckling was detected at 85 °C, with no actuation above this temperature. The operational interval is however wide enough for most of the possible applications. In the case of cantilever switch, the membrane is always stress-free, and the actuation voltage variation is much more limited, being of the order of 11% in the interval 25–100 °C, and the operational interval
much wider. Other temperature related effects are responsible in this case for actuation voltage variation, such as elastic modulus variation, and charging. The complete restoring of the initial actuation voltages after a week proves that creep does not play a role in the variation of cantilever actuation voltage. A design which minimizes charging can further reduce the temperature dependence for the cantilever typology, and considerably broaden its operational interval. More in general, different switch typologies, not investigated in this paper but especially studied to reduce temperature stress variation both at design and at material level, can also reduce the temperature dependence of the switch parameters in clamped–clamped structures. Acknowledgment This work has been partially supported by ESA/ESTEC within the contract ESA ITT AO/1-5288/06/NL/GLC ‘‘High Reliability MEMS Redundancy Switch’’. References [1] Papaioannou G, Exarchos M-N, Theonas V, Wang G, Papapolymerou J. Temperature study of the dielectric polarization effects on capacitive RFMEMS switches. IEEE Trans Microw Theory Technol 2005;53:3467–73. [2] Yuan X, Peng Z, Hwang JCM, Forehand D, Goldsmith CL. Acceleration of dielectric charging in RF-MEMS capacitive switches. IEEE Trans Device Mater Rel 2006;6:556–63. [3] Zhu Y, Espinosa H. Effect of temperature on capacitive RF-MEMS switch performance – a coupled-field analysis. J Micromech Microeng 2004;14: 1270–9. [4] Mulloni V, Giacomozzi F, Margesin B. Controlling stress and stress gradient during the release process in gold suspended micro-structures. Sens Actuators A 2010;162:93–9. [5] Tilmans HAC. MEMS components for wireless communications. In: EUROSENSORS XVI Prague Czech Republic 15–18, September 2002, p. 1–34. [6] Chen K-S. Techniques in residual stress measurement for MEMS and their applications. In: MEMS/NEMS Handbook – Techniques and Applications, vol. 1. Springer; 2007. p. 1252–1320. [7] Rebeiz GM. RF-MEMS: theory, design and technology. New York: Wiley; 2003. [8] Yan X, Brown WL, Li Y, Papapolymerou J, Palego C, Hwang JCM, et al. Anelastic stress relaxation in gold films and its impact on restoring forces in MEMS devices. JMEMS 2009;18:570–6. [9] Giacomozzi F, Mulloni V, Colpo S, Iannacci J, Margesin B, Faes A. A flexible fabrication process for RF-MEMS devices. Romanian J Inf Sci Technol 2011;14:259–68. [10] Rottemberg A, Jansen H, Fiorini P, De Raedt W, Tilmans H. Novel RF-MEMS capacitive switching structures. In: Proc Eumic 2002, September 24–26, Milan, Italy; 2002. p. 809–12. [11] Bartolucci G, Marcelli R, Catoni S, Margesin B, Giacomozzi F, Mulloni V, et al. An equivalent circuital model for shunt connected coplanar RF-MEMS switches. J Appl Phys 2008;104:84514. [12] Brusa E, Munteanu MG. Role of the electro-thermo-mechanical multiple coupling on the operation of RF-microswitch. Microsyst Technol 2012;18:983–95. [13] Lafontan X, Le Touze C, Wenk B, Kolesnik I, Pressecq F, Perez G, Nicot J-M, Dardalhon M, Rigo S. Environmental test bench for reliability studies:
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influence of the temperature on RF switches with metallic membranes. DTIP2002 SPIE, vol. 4755; 2002. p. 624–33. [14] Palego C, Solazzi F, Halder S, Hwang JCM, Farinelli P, Sorrentino R, Faes A, Mulloni V, Margesin B. Effect of substrate on temperature range and power capacity of RF-MEMS capacitive switches in IEEE EuMC 2010; September 2010. p. 505–8. [15] Solazzi F, Tazzoli A, Farinelli P, Faes A, Mulloni V, Margesin B, et al. Design and characterization of an active recovering mechanism for high-performance RFMEMS redundancy switches. J Microsyst Wireless Tech 2011:539–46.
[16] Palego C, Deng J, Peng Z, Halder S, Hwang JCM, Forehand DI, et al. Robustness of RF-MEMS capacitive switches with molybdenum membranes. IEEE Trans Microw Theory Technol 2009;57:3262–9. [17] Neighbours JR, Alers GA. Elastic constants of silver and gold. Phys Rev 1958;111:707–12. [18] Beake BD, Smith JF. High-temperature nanoindentation testing of fused silica and other materials. Philos Mag A 2002;82:2179–86.
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Contents lists available at SciVerse ScienceDirect
Microelectronics Reliability journal homepage: www.elsevier.com/locate/microrel
Influence of temperature on the actuation voltage of RF-MEMS switches V. Mulloni ⇑, F. Solazzi, F. Ficorella, A. Collini, B. Margesin Fondazione B. Kessler, Center for Materials and Microsystems, Via Sommarive 18, I-38123 Trento, Italy
a r t i c l e
i n f o
Article history: Received 7 September 2012 Received in revised form 17 January 2013 Accepted 17 January 2013 Available online xxxx
a b s t r a c t Most of the actual applications for RF-MEMS switch require high reliability, but consolidated qualification procedures are still lacking. This paper focuses, in particular, on the role of temperature on the switch reliability from a mechanical point of view, showing how this depends on the switch architecture and membrane material. Double clamped switches are sensitive to buckling, and this is the factor limiting their operational temperature, even though the range exploitable can be wide enough for many applications. Residual stress and thermal expansion coefficient of the mobile membrane are the most important parameters to understand and control this phenomenon. Cantilever switches are less influenced by the temperature in their performances, and have a much wider operational range. Other temperature-related factors are affecting the switch reliability in this case, such as elastic modulus variation, dielectric charging effects and creep. Ó 2013 Elsevier Ltd. All rights reserved.
1. Introduction In the last years RF-MEMS switches have been developed by several research groups and companies, demonstrating outstanding performances in terms of RF-characteristics in a broad range of frequencies. Nonetheless RF-MEMS switch reliability is still an open issue which so far prevented an extensive use of these devices. In applications where temperature variations are important, such us in satellite communications, or in the packaging process itself, the effect of temperature on device performances is of great importance for evaluating device reliability, and usually a defined operational temperature interval is explicitly required to qualify a defined switch technology. Temperature can affect switch characteristics in several ways. Electrical polarization mechanisms due to charging, such us free charge redistribution, charge injection and charge dissipation are temperature dependent [1,2], but the mechanical properties of the switching membrane are also strongly affected by the temperature [3], and the combined effects of all these physical variations can couple and superimpose increasing the complexity of reliability testing and analysis. In this view, the data presented in this paper focuses only on the mechanical modifications induced by the temperature on the MEMS mobile structure, with the purpose to separate these effects from other types of temperature induced reliability issues when analyzing switch reliability data and trying to design accelerating stress tests. Usually, after the release process, the mobile part of MEMS switch is deformed to some extent in the out-of-plane direction, due to a complicated interplay of its residual stress and stress ⇑ Corresponding author. E-mail address: [email protected] (V. Mulloni).
gradient and the modifications induced by the release process itself [4]. This out-of plane profile may sensibly change with temperature, and needless to say, the actuation properties of the device are strongly influenced by these variations. The key mechanical parameter to understand these modifications is however not the same in all type of switches, and in particular, is quite different in the case of double-clamped or single clamped switches, which are the most common switch typologies. In this paper the temperature-induced actuation voltage variations in the temperature interval 25–100 °C are measured, interpreted and discussed for a double-clamped and a cantilever switch. Even though the switching membrane are made of the same material, the operational range and the actuation voltage variations are very different for the two typologies. The main reasons for this differences are described, resulting in a stress dominated behavior in the case of double-clamped switch, while other factors are affecting the behavior of cantilever switch. 2. Temperature related mechanical variations 2.1. Double-clamped structures In the case of a switch membrane clamped at both ends, the key mechanical parameter influenced by the temperature is the residual stress of the membrane. The usual analytical formula reported for the actuation voltage Vpi of a switch is [5]:
sffiffiffiffiffiffiffiffiffiffiffiffiffi 3 8kd0 V pi ¼ 27e0 A
ð1Þ
where d0 is the zero-voltage gap spacing, A the electrode area, e0 is the vacuum dielectric permittivity and k is the spring constant.
0026-2714/$ - see front matter Ó 2013 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.microrel.2013.01.007
Please cite this article in press as: Mulloni V et al. Influence of temperature on the actuation voltage of RF-MEMS switches. Microelectron Reliab (2013), http://dx.doi.org/10.1016/j.microrel.2013.01.007
2
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While d0, A, and e0 can be considered temperature independent within a wide range, the spring constant is strongly temperature dependent [6]. In a double-clamped structure, k is determined by two main contributions [7]: one is given by the elastic response to an applied force and is proportional to the Young’s modulus of the structure, while the second part is determined by the residual stress of the membrane. In the most common case, the switch substrate is silicon or quartz, and the mobile membrane is made of metal. In this case, being the membrane anchored to the substrate, the different thermal expansion coefficients produce an additional tension in the membrane when the temperature is lowered with respect to the ambient temperature, and a compression when the temperature is raised. The membrane itself usually possesses an intrinsic stress that may strongly vary with processing conditions, but it is normally tensile. The switch actuation voltage decreases with decreasing the tensile membrane stress, that is it decreases with increasing temperature until it reaches a minimum value, while its shape does not change and the gap remains constant. After this minimum value, a discontinuity appears because the membrane buckles and the switch can no longer properly actuate. This behavior can be represented by the following equation:
V 2pi / k ¼ kelastic ðEÞ þ B½rresidual
Tamb
Eðametal asubstrate ÞDT
ð2Þ
where DT is the temperature difference with respect to ambient, which is taken as reference temperature, E is the effective Young’s modulus of the metal, ametal and asubstrate are the thermal expansion coefficients of the metal and the substrate respectively, and B is a convenient geometric factor which depends on the switch geometry and dimensions. The minimum actuation voltage is reached when the residual stress of the membrane rresidual_Tamb is perfectly compensated by the last term of Eq. (2), which is the thermal stress. At the temperature in which this occurs the membrane is stress free, and its actuation voltage is determined only by its geometry and by the elastic modulus of the metal. However, the stress interval between zero stress and buckling is very narrow [6], as can be estimated by the Eulero’s formula. For practical purposes it is correct to assume that as soon as the stress turns compressive, the membrane buckles. For double-clamped switches then, the range of operational temperatures is then defined by the maximum acceptable actuation voltage (lowest temperature) and by the buckling temperature (highest temperature). Other factors can play a role, such as the variation of Young’s modulus with temperature, which decreases with increasing temperature, adding its effect to that of thermal stress, and charging phenomena, that may become faster at higher temperature [2]. However, they are estimated to be much less relevant with respect to the stress contribution. 2.2. Single-clamped structures Single clamped switches are of cantilever type and in first approximation can be considered stress free, because they can expand or contract in response to temperature variations. This means that the term within square brackets in Eq. (2) vanishes and k = kelastic(E). Ruling out the major temperature-related source of Vpi variation, other effects such as charging, Young’s modulus variation and, when the temperature is high enough, plastic relaxation and creep phenomena may become visible. The role of elastic modulus variation can be understood considering that in first approximation k depends linearly on E [7], and then, in force of Eq. (1), Vpi depends on the square root of E. Moreover, even though for this switch typology the average biaxial stress is virtually zero, many cantilever structures possess a relevant stress gradient which results in a marked non-planarity of the suspended membrane. This gradient develops during the release phase [4] and could be temperature dependent. In this case
the actuation voltage can be affected by the gradient variation resulting in a change in the average actuation gap. This effect, when present, can be potentially important, given the marked dependence of Vi from the average actuation gap d0 in Eq. (1). As reported before, all the mentioned effects are estimated to be much less severe than the stress variation in the case of clamped–clamped switches, and then the operational temperature range of cantilever structures is predicted to be much wider than that of clamped–clamped membranes. In particular the maximum operational temperature, when high enough, is likely to be determined by creep [8], which is strongly temperature dependent and may become sensible especially after prolonged actuation. The distinct contributions may however widely vary with switch design and materials.
3. Experimental The MEMS switches investigated were fabricated with an eightmask process at FBK. A detailed process description is discussed in [9], and can be briefly summarized as follows: high resistivity 525 lm thick and 4 in. wide silicon wafers were used as substrate. Actuation electrodes are made in lightly doped polysilicon, while the underpass is realized with a 650 nm-thick sputtered aluminum layer. The central underpass line is covered with 100 nm of PECVD oxide and then with 150 nm of a floating potential metal, which is made by e-gun evaporated gold. The mobile membrane is made by a 2 lm thick electrodeposited gold layer, suspended over an air gap of 2.7 lm, obtained by burning in oxygen plasma a sacrificial photoresist spacer. A second and thicker (about 3.5 lm) electrodeposited gold layer is present in the coplanar waveguide and in the anchoring region. In the case of the double clamped capacitive switch, the role of the floating metal electrode is to ensure a good electrical contact with the mobile membrane and the dielectric allowing a very high and predictable Con/Coff ratio [10,11], which is around 200 in the selected design, reported in Fig. 1a. In addition, the dielectric is removed from the actuation pads so to minimize charging phenomena, while mechanical stoppers, placed within but not in contact with the electrode area, prevent the contact between the downstate bridge and the actuation pads. For what concerns the cantilever switch, a picture of the selected device is reported in Fig. 1b. In this case the switch is ohmic, non mechanical stopper is present and the dielectric layer has not been removed above the electrode .This means that no particular design strategy has been taken to eliminate or reduce charging effects. Six polysilicon dimples covered by the underpass and by the 150 nm thick gold layer stabilize the gold–gold contact in the switching region. The switching membrane is made again with a 2 lm thick electrodeposited gold layer, while the 3.5 lm thick gold layer is present on part of the membrane to reinforce it, in order to compensate for the lower restoring force of cantilever-type switches. Surface topographies and vertical profiles were taken with a Zygo optical profiler, with a 10 magnification. The vertical resolution is about 1 nm, and the lateral resolution is 1 lm. The profiles at different temperature where taken heating the wafer at 120 °C and then monitoring the substrate temperature in the cooling phase during the measurement. It is then possible that the real temperature on the switch was few degrees above or below the measured one. Electrical measurements have been performed with an automated measurement systems made by an Accretech UF200A Wafer Prober, with a temperature controlled chuck, an Agilent 4284A Precision LCR Meter, an Agilent E5270B Measure Mainframe equipped with six Precision Source Monitor Units and an Agilent
Please cite this article in press as: Mulloni V et al. Influence of temperature on the actuation voltage of RF-MEMS switches. Microelectron Reliab (2013), http://dx.doi.org/10.1016/j.microrel.2013.01.007
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Fig. 1. Pictures of the RF-switches investigated: (a) Clamped–clamped switch. (b) Cantilever switch.
4. Results and discussion 4.1. Clamped–clamped switch The actuation voltages as a function of the chuck temperature are reported in Fig. 2, for seven identical switches distributed on different parts of the same wafer. For all the switches the actuation voltage decreases markedly with the temperature up to 70 °C, going from 42 ± 4 V down to 24 ± 2.5 V with a common trend that is similar to what expected, with the only exception of sample 1 which deviates from the expected trend already at 70 °C. The actuation voltages at 85 °C show an erratic behavior and none of the switches actuates at 100 °C. The dashed curves are the best fit obtained for each switch fitting linearly the square of Vpi as a function of the temperature in the range 20–70 °C, according to Eq. (2). It is worth noticing that the different fitting curves converge around
60
Sample 1 Sample 3 Sample 5 Sample 7
50
Actuation Voltage (V)
B2201A Low-Leakage Switching Matrix, which are all managed by an in-house developed software framework. In the case of the double-clamped capacitive switch a C–V measurement was performed with the use of the Precision LCR Meter. A bias voltage was applied between the actuation pads and the mobile bridge while the capacitance between the underpass and the bridge was measured with a signal at 1 MHz frequency and 30 mV of amplitude. The Vpi was detected when the capacitance value instantly increases from Coff to Con value. In the case of the cantilever resistive switch a R–V measurements was performed. A bias voltage was applied between the actuation pads and the mobile bridge, while the resistance between the underpass and the bridge was measured. The Vpi was detected when the bridge is connected to the underpass, allowing the current to flow and then instantly switching from an open circuit off state to a low series resistance on state. For each device and temperature, four consecutive sweeps have been generated starting from 0 V to 75 V, then from 75 V to 0 V, then from 0 V to 75 V and finally from 75 V to 0 V, without voltage interruption and with a voltage step of 0.5 V. The time required for the measurement of a single device at a given temperature was about 5 min. A recovery time of about 24 h at 25 °C was imposed between two measurements with different chuck temperatures. Temperature control was performed at wafer level by the Prober itself, with a fluctuation which is less than 1 °C. Before the measurement starts, the wafer was warmed-up at the selected temperature and allowed to thermalize, then all the devices on the wafer were sequentially measured.
Sample 2 Sample 4 Sample 6
40
30
20
10 20
30
40
50
60
70
80
90
Temperature (oC) Fig. 2. Actuation voltage as a function of the temperature for seven clamped– clamped switches on the same wafer.
80 °C, with the only exception of Sample 1, which has however been fitted only with the first three experimental points. While it is reasonable to assume that the membrane shape does not change up to 70 °C, it is straightforward to associate the unpredictable behavior at 85 °C with the emerging of buckling, and the convergence point around 80 °C can be reasonably interpreted as the temperature where the minimum actuation voltage can be reached. The data in Fig. 2 correlate well with the longitudinal profiles of the gold membrane reported in Fig. 3b, measured along the dashed lines of the surface topography of Fig. 3a. The profile at 25 °C shows that the membrane is not perfectly planar even at room temperature. The out of plane features are mainly due to the incomplete planarization of the sacrificial photoresist spacer which accommodates only partially the underneath vertical steps due to the presence of the actuation electrodes and underpass line. The underpass is actually higher than the electrodes and the stoppers due to the presence of the floating metal layer. The interesting feature visible in Fig. 3b is however the stability of the longitudinal profile in the range 20–65 °C and then the progressive enhancement of the central part of the membrane at the temperature of 85 °C and even more at 100 °C. The actuation voltages at 85 °C are in any case higher that those predicted by the fitting curves and this can be explained considering the increased effective gap as a consequence of the membrane
Please cite this article in press as: Mulloni V et al. Influence of temperature on the actuation voltage of RF-MEMS switches. Microelectron Reliab (2013), http://dx.doi.org/10.1016/j.microrel.2013.01.007
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Fig. 3. (a) Surface map and (b) longitudinal profiles taken along the dashed arrow in (a) for one of the clamped–clamped switches investigated.
Actuation Voltage (V)
70
60
Sample 1 Sample 3 Sample 5 Sample 7
50
Sample 2 Sample 4 Sample 6 Sample 8
40
20
30
40
50
60
70
80
90
100
25 bis
25 ter
Temperature (oC) Fig. 4. Actuation voltage as a function of the temperature for eight cantilever switches on the same wafer. Data labeled 25 bis was measured at 25 °C just after the measurement at 100 °C. Data labeled 25 ter was measured at 25 °C after a week.
bending in the upper direction. At 100 °C the bending is so pronounced that the switch cannot properly actuate because the central part of the membrane cannot touch the floating electrode underneath, being blocked by the stoppers placed on the lateral actuation electrodes. It should be noted that we cannot exclude, as theoretical possibility, a switch bending in the downward direction as the effect of buckling [12], and in this case a strong reduction of actuation voltage or stiction should be the consequence. Since this does not happen for any of the switches investigated, it is possible that the switch geometry or the presence of a stress gradient in the membrane layer forces the membrane to bend in the upward direction in the unactuated position. In fact, as will be shown further in this section, the cantilever switches fabricated in the same wafer show a marked upward bending, as reported in Fig. 5b, which indicates a markedly positive stress gradient, and this can explain the observed upward buckling. We know this gradient develops during the release phase [4], and we found it quite reproducible within the adopted switch fabrication technology from batch to batch. Some numerical analysis could give a better explanation of this experimental phenomenon, but it is beyond the purpose of this paper which reports mainly experimental data and could be a possible topic for future work. From Fig. 2 we can assume that 80 °C is approximately the temperature in which the zero stress is reached, and we can calculate
the membrane stress simply assuming that the part of Eq. (2) enclosed within square brackets is equal to zero. In the case of silicon substrate and gold membrane, ametal and asubstrate are equal to 2.3 106 K1 and 14.2 106 K1, respectively [13], while the Young’s modulus of gold is around 75 GPa. Being the temperature difference of 55 °C, the result for the residual stress at 25 °C is approximately 49 MPa. This value is very close to the stress values obtained from simulations on similar structures [14,15]. We can assume that highest operational temperature for this type of switch is then around 80 °C, and we can set the lower bound at least 20 °C for all the switches investigated, if we set an acceptable maximum actuation voltage at 60 V, using a prediction made with the curves in Fig. 2. The effective operational range is then sufficient for most of the possible application, even though the Vpi variations are quite large within this range. It is important to note that the operational range may strongly vary not only with the value of room temperature stress, but also with the type of metal used for the membrane. Residual stress, being mostly a result of processing conditions, is quite hard to control during manufacturing and, while a high stress value can raise the maximum operational temperature, it can seriously worsen other reliability problems. The choice of a metal with a low thermal expansion coefficient can be instead a big advantage [16]. For example, with a mobile membrane made of aluminum, but with the same stress value of those previously presented, the maximum operational temperature would be only 59 °C, being its Young’s modulus close to that of gold (70 GPa) but its ametal is, almost double (23 106 K1) [3], while switches made with molybdenum membrane can operate up to 150 °C [16], due to the very low thermal expansion coefficient of molybdenum (4.8 106 K1). 4.2. Cantilever switch The actuation voltages for positive bias as a function of the temperature are reported in Fig. 4 for eight cantilever switches with the design of Fig. 1b distributed on different parts of the same wafer. The dashed lines in Fig. 4 are only a guide for the eyes. The Vpi’s decrease with the temperature, but the decrease from 25 to 100 °C is limited to about 11% in all cases, and all the switches actuate up to 100 °C. The switch topography together with the profiles taken at different temperatures for one of the samples investigated are reported in Fig. 5 where two things appear clearly: first, the switch membrane bends upwards, due to an evident positive stress gradient, with a height difference from anchor to tip of 3.5 lm; second, the temperature has no visible effect on this bending within the range investigated. This rules out the gradient variation phenom-
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tip
20 oC 85 oC 100 oC
6
anchor
Height (µm)
8
4
(a)
(b)
2 0.05
0.10
0.15
0.20
0.25
Length (mm) Fig. 5. Surface map and (b) longitudinal profiles taken along the blue dashed arrow in (a) for one of the cantilever switches investigated.
ena to explain the actuation voltage decrease. A partial role in this decrease can be attributed to variations of gold elastic modulus. From the data in Refs. [17,18] this variation can be estimated of the order of 4–5% in the temperature interval 25–100 °C and then the Vpi reduction should be no more than 3%. This implies that there is some other source of temperature-induced actuation voltage change. Charging is the most likely candidate, since the switch has not been designed to minimize it, and because charging phenomena are known to accelerate with temperature [2]. It is then possible that the switch is not fully discharged when measured again at a different temperature. An evidence of the presence of charging can be given by the comparison of positive and negative actuation voltages obtained in the bipolar scan. In fact, some differences have been detected, in the range of 0.5–2 V, indicating that some charging is present, but also that the values are close to the experimental error, being the sampling interval of 0.5 V, and no clear temperature trend was detected. Furthermore, this test evidences the presence of charging but does not gives us the absolute value of this shift. Another possibility to verify the presence of charging is to measure again the Vpi at 25 °C after the measurement at 100 °C. The data obtained in this way are shown in Fig. 4 after the data at 100 °C corresponding to the label 25 bis. As can be seen, the recovery towards the initial position is minimal. This means that the measured variations are either not due only to temperature (charging or other possible but unknown effects), or are not fully reversible (creep). Creep cannot be completely excluded from data in Fig. 5b, because it can be induced by the stress of actuation, which was not been applied when the data in Fig. 5b was measured. To separate this two effects a third measurement at 25 °C was performed after a week, which is a time surely sufficient to dissipate any residual charging. The corresponding data is labeled 25 ter in Fig. 4. As can be seen, the recovery is almost complete, indicating that the measured variations are likely due to charging and no irreversible effect took place.
5. Conclusion The variation of actuation voltage with temperature was measured and discussed for two switch typologies. For clamped– clamped switches, the variations are more than 40% in the temperature interval 25–70 °C. In this case the membrane stress variation is the dominant effect, and buckling was detected at 85 °C, with no actuation above this temperature. The operational interval is however wide enough for most of the possible applications. In the case of cantilever switch, the membrane is always stress-free, and the actuation voltage variation is much more limited, being of the order of 11% in the interval 25–100 °C, and the operational interval
much wider. Other temperature related effects are responsible in this case for actuation voltage variation, such as elastic modulus variation, and charging. The complete restoring of the initial actuation voltages after a week proves that creep does not play a role in the variation of cantilever actuation voltage. A design which minimizes charging can further reduce the temperature dependence for the cantilever typology, and considerably broaden its operational interval. More in general, different switch typologies, not investigated in this paper but especially studied to reduce temperature stress variation both at design and at material level, can also reduce the temperature dependence of the switch parameters in clamped–clamped structures. Acknowledgment This work has been partially supported by ESA/ESTEC within the contract ESA ITT AO/1-5288/06/NL/GLC ‘‘High Reliability MEMS Redundancy Switch’’. References [1] Papaioannou G, Exarchos M-N, Theonas V, Wang G, Papapolymerou J. Temperature study of the dielectric polarization effects on capacitive RFMEMS switches. IEEE Trans Microw Theory Technol 2005;53:3467–73. [2] Yuan X, Peng Z, Hwang JCM, Forehand D, Goldsmith CL. Acceleration of dielectric charging in RF-MEMS capacitive switches. IEEE Trans Device Mater Rel 2006;6:556–63. [3] Zhu Y, Espinosa H. Effect of temperature on capacitive RF-MEMS switch performance – a coupled-field analysis. J Micromech Microeng 2004;14: 1270–9. [4] Mulloni V, Giacomozzi F, Margesin B. Controlling stress and stress gradient during the release process in gold suspended micro-structures. Sens Actuators A 2010;162:93–9. [5] Tilmans HAC. MEMS components for wireless communications. In: EUROSENSORS XVI Prague Czech Republic 15–18, September 2002, p. 1–34. [6] Chen K-S. Techniques in residual stress measurement for MEMS and their applications. In: MEMS/NEMS Handbook – Techniques and Applications, vol. 1. Springer; 2007. p. 1252–1320. [7] Rebeiz GM. RF-MEMS: theory, design and technology. New York: Wiley; 2003. [8] Yan X, Brown WL, Li Y, Papapolymerou J, Palego C, Hwang JCM, et al. Anelastic stress relaxation in gold films and its impact on restoring forces in MEMS devices. JMEMS 2009;18:570–6. [9] Giacomozzi F, Mulloni V, Colpo S, Iannacci J, Margesin B, Faes A. A flexible fabrication process for RF-MEMS devices. Romanian J Inf Sci Technol 2011;14:259–68. [10] Rottemberg A, Jansen H, Fiorini P, De Raedt W, Tilmans H. Novel RF-MEMS capacitive switching structures. In: Proc Eumic 2002, September 24–26, Milan, Italy; 2002. p. 809–12. [11] Bartolucci G, Marcelli R, Catoni S, Margesin B, Giacomozzi F, Mulloni V, et al. An equivalent circuital model for shunt connected coplanar RF-MEMS switches. J Appl Phys 2008;104:84514. [12] Brusa E, Munteanu MG. Role of the electro-thermo-mechanical multiple coupling on the operation of RF-microswitch. Microsyst Technol 2012;18:983–95. [13] Lafontan X, Le Touze C, Wenk B, Kolesnik I, Pressecq F, Perez G, Nicot J-M, Dardalhon M, Rigo S. Environmental test bench for reliability studies:
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[16] Palego C, Deng J, Peng Z, Halder S, Hwang JCM, Forehand DI, et al. Robustness of RF-MEMS capacitive switches with molybdenum membranes. IEEE Trans Microw Theory Technol 2009;57:3262–9. [17] Neighbours JR, Alers GA. Elastic constants of silver and gold. Phys Rev 1958;111:707–12. [18] Beake BD, Smith JF. High-temperature nanoindentation testing of fused silica and other materials. Philos Mag A 2002;82:2179–86.
Please cite this article in press as: Mulloni V et al. Influence of temperature on the actuation voltage of RF-MEMS switches. Microelectron Reliab (2013), http://dx.doi.org/10.1016/j.microrel.2013.01.007