The sensitivity of an innovative acoustic sensor composed of four hot-wires is analyzed. An analytical model is presented that describes both the air flow and the temperature distribution in and around the probe. The presence of the chip surface in the vicinity of the wires influences the acoustic flow, while it also affects the temperature distribution. Both effects result into a specific angular dependence of the sensor sensitivity.\ud
Acoustic measurements are compared with theory, showing good correspondence

Elwenspoek, M.

Honschoten, J.W. van

Wiegerink, R.J.

Yntema, D.R.

English

The sensitivity of an innovative acoustic sensor composed of four hot-wires is analyzed. An analytical model is presented that describes both the air flow and the temperature distribution in and around the probe. The presence of the chip surface in the vicinity of the wires influences the acoustic flow, while it also affects the temperature distribution. Both effects result into a specific angular dependence of the sensor sensitivity. Acoustic measurements are compared with theory, showing good correspondence

van Honschoten, J.W.

Yntema, Doekle Reinder

Wiegerink, Remco J.

Elwenspoek, Michael Curt

University of Twente Research Information

MODELING AND CHARACTERIZATION OF THE SENSITIVITY OF A HOT -WIRE PARTICLE VELOCITY SENSOR  J.W. van Honschoten, D.R. Yntema, R.J. Wiegerink, M. Elwenspoek1 1MESA+ Research Institute, University of Twente,  PO 217, 7500 AE Enschede, The Netherlands j.w.vanhonschoten@utwente.nl Abstract The sensitivity of an innovative micromachined acoustic sensor consisting of four hot-wires is analyzed theoretically and experimentally. An analytical model is presented that describes both the air flow and the temperature profile in and around the probe. The presence of the chip surface in the vicinity of the wires influences the acoustic flow, while also affecting the temperature distribution in the probe. Both effects result into a specific angular dependence of the sensor sensitivity. Acoustic flow measurements are compared with the theory and with numerical simulations on the device,showing good qualitative and quantitative correspondence. Key Words: particle velocity sensor, acoustics, flow sensors, MEMS-modelling I  INTRODUCTION For a complete characterization of a three-dimensional sound field, both the acoustic particle velocity and the pressure have to be determined. By means of a particle velocity sensor complementary to the conventional pressure microphone, a complete determination of the three-dimensional acoustic field, including intensity streamlines and different source contributions becomes possible. To this purpose a micromachined particle velocity sensor, composed of two parallel heated wires, of about 1 mm length and spaced about 300 µm apart, has been developed [1,2,3,4]. When subject to an acoustic flow, forces convection leads to a small temperature difference between the wires, which is ameasure of the particle velocity. To exploit fully its abilities to measure the three-dimensional particle velocity vector, we designed a new device in which four hot-wires in perpendicular directions are integrated in one probe [5] (Fig.1).  For a very accurate determination of the direction of the involved particle velocities, a good understanding of the directional sensitivity of the probe is needd. However, the current designs show a deviation in their directional sensitivity pattern from the ideal ‘figure-of eight’ response [2,6] that would be the case for free standing wires. Up to now, this deviation has not been fully understood. The purpose of this paper is to analyze this effect and model the sensitivity for the four wire probe, with the sensor wires placed within a square orifice in the chip surface. In the analysis we take into account all physical aspects that play a role: both the acoustic flow profile around the wires, that is disturbed due to the presence of the printed circuit board acting as a disturbance for theflow, and the local temperature distribution close to the heaters. The chip surface acts as a heat sink that affec s the heat transport due to convection and consequently influences the sensor response to particle velocities. We present an analytical model for the directional sensitivity of this novel device, and compare this w th experiments and numerical calculations on the flow and temperature distribution in the probe.   Figure 1. (upper:) The four wire particle velocity probe, containing four 1 mm long and 400 µm wide sensor wires suspending the rectangular orifice in the chip surface. (lower:) schematic picture of the probe. II  THEORY Ideally, the two parallel heated wires of a particle velocity sensor exhibit a cos(θ)-dependent response to a fluid flow imposed under an angle θ with the plane of the wires, since the sensor wires detect the velocity component directed in the plane of the wires only. This cos(θ) angular dependence on the incoming flow is sometimes referred to as a ‘figure-of-eight’ response [2,6]. However, in practice free standing wires without any connections to the probe do not exist, and as aconsequence of the surrounding material the angular dependence of the sensitivity is significantly influenced. For the probe as shown in figure 1, where the thin wires (1 mm long silicon nitride beams of 400 µm thickness with 100 nm thick platinum layers on top) are placed within a square orifice in a 380 µmthick silicon support, measurements show a significant constant deviation from the figure-of-eight pattern. Is is observed that the measured angle deviates from the angle θ of the incoming flow; but also that this angular shift of approximately 18o appears to be θ-independent as well, which results into a constant rotation of theentire ‘figure-of-eight’ curve of about 18o in the polar plane. Two possible causes for the observed effect can be propounded. First, the presence of the chip surface in the close proximity of the wires disturbes the air flow, thereby altering the direction of the local particle velocity. Second, the silicon chip acts as a ‘heat sink’ resulting in a preferred direction of the heat flow generated by the sensor wires, to the silicon. Consequently, the temperature distribution around the wires is not completely symmetric anymore as it would be for wires in free space. II.1  THE TEMPERATURE PROFILE To analyze the response of the sensor due to an imposed acoustic flow, the temperature profile around the heated wires has to be calculated. At first, the static temperature distribution in the device is calculated, when no gas flow is present. It has been proven [7,8]that an acoustic wave gives only a small perturbation to this temperature profile, and also that the output signal of the sensor is linearly proportional to the local temperature gradient at the place of the sensor wires. Therefore, the static temperature distribution only provides sufficient information to deduce the sensitivity properties we are interested in in thispaper, in particular  the directionality (the polar dependce) of the sensitivity of the device.  In the absence of a moving gas, the temperature distribution around the wires follows from the stationary heat equation ( ) QTk =∇∇−      (1) where k = k(T) is the thermal conductivity of air, that is in principle a function of temperature, and Q the heat produced by the heater per unit volume per unit time. We now solve the static temperature T(x,z) for the geometry as shown in Fig. 2. Taking into account the appropriate boundary conditions for T and the heat flux at x = ± lx, we can obtain an expression for the temperature in the device as an expansion in harmonics in normalized variables ξ = x/lx, ζ = z/lx. The resulting expansion converges quickly and can be easily calculated numerically. For a specific set of parameters the temperature distribution is shown in figure 3, also compared with numerical simulations described in section III.     lx    D   d          ↑         z           d                               x →   Figure 2. Cross-sectional view of the model geometry for the calculation of the temperature distribution. The thickness of the substrate (the chip) is d; its full width is 2D, the horizontal distance between the wires is also d.  Figure 3. Temperature distribution in the channel along the normalised x-axes, according to the calculated expansion for T with ξ0=0.2, lx = 1 mm and ly=1.5 mm. For comparison, results from numerical simulations (section III) are also shown. The temperatures are normalised by T0, the temperature at the place of the heaters. The thus obtained static temperature profile is the required starting point to find the sensitivity S of the device and its response to an acoustic gas flow. If due to an acoustic wave the gas close to the heaters has a velocity v, the temperature around the wires is described by the full heat equation ( ) ( ) QTkTvTc tp =∇∇−∇+∂ρ    (5) where cp is the gas heat capacity at constant pressure. The convective term is responsible for the difference in wire temperatures. For small v, the solution of this equation is formed by the previously found temperature plus a small correction δT proportional to Tv ∇⋅ , with T the previously obtained static temperature [7,9]. The sensitivity S is proportional to δT and we obtain an expression for the angular dependence S(θ), with θ the angle of the incident flow. It is seen that the representation of S(θ) in the polar plane, the polar figure, is shifted with respect to the ‘ideal’ figure-of-eight of the sensitivity of free standing wires, byan angle ∆, determined by the ratio of the x-and z-derivatives of T at the wires (Fig.4). The calculated shift ∆, due to this temperature effect, as a function of the gap distance is shown in Fig. 7.   0.2  0.4  0.6  0.8  1302106024090270120300150330180 0  experimenttheorynum.calc. Figure 4. Polar figure of the sensitivity of the probe at 600 Hz, compared with the model and the numerical calculations. An ‘ideal’ figure of eight would have its maximum at 45o; the measured angular shift ∆ is 18o. II.2 THE AIR FLOW AROUND THE PROBE Due to the presence of the chip surface, the fluid lowitself is influenced as well, and the local angle q will in general not be equal to the angle of incidence of the acoustic wave far away from the probe. To analyse therefore the air flow around the probe, one should solve the full Navier-Stokes equations for the three-dimensional probe geometry to obtain the complete particle velocity field around the probe. For the relatively complicated geometry of the real probe as depicted in figure 1, an exact analytic solution canot be found. Therefore, we model the sensor chip by a pair of long and thin parallel plates, of which thecross section is shown by figure 2. The problem becomes then two-dimensional. It has been shown [8, 10] thatfor propagating acoustic waves, the fluid can be considered as incompressible, so that the Navier-Stokes equations yield vpvvtv 21)( ∇+∇−=∇⋅+∂∂ νρ  (9) with p the pressure and  ν the kinematic viscosity. It was shown before [8,10], that for the current values of v, the nonlinear term in the equation can be neglected. Using elliptic coordinates [11], the equation can be solved analytically for thin, infinitely long, plates of width D-lx (d << ( D-lx) ) For a specific value of θ, the streamline pattern is seen in Fig. 5. From the obtained streamline pattern, we can determine the local velocity of the air at the place of the wires, and determine the angular shift ∆. The values of ∆ due to this flow profile around the probe, are also depicted in Fig.7.  Figure 5. Calculated streamline pattern in the normalized x-z-plane (see the cross section of Fig.2) for an incident flow with an angle of incidence θ = π/8, and a normalized gap distance of the probe equal to 2. III  NUMERICAL CALCULATIONS We used a finite-elements numerical solver, CFD RC ®, to analyze the temperature distribution and the air flow in and around the three-dimensional probe geometry. Both the Navier-Stokes equations and the heat diffusion equations were solved numerically by constructing a fine mesh of discrete small control volumes in the solution space, defining the appropriate boundary and initial conditions, and solving the equations for each cell. A plane propagating wave of incident angle θ was defined as entering the outer boundaries, and θ was varied in the simulations. The dissipated power per wire was set to be 20 mW. Devices with varying gap distances, ranging from 0.2 mm <lx < 1.8 mm, were investigated. Additionally, the heat                                            x (mm)                                            x (mm)   T/T0 1.0 0.9 0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0   Figure 6. Numerical simulations on the temperature profile in the probe. Shown are cross sections of the chip, on which two thin sheets of material are deposited, displaying the isotherms (normalised by T0 at the wires). Figure left: The material has a heat conductivity kair = 0.04 W/m K (corresponding to the wood that was used in the experiments); right: with kSi = 130 W/m K. conductivity of the chip material was varied, ranging from the heat conductivity of air, kair = 0.0263 W/m K, to that of silicon, kSi = 130 W/m K, in five linear steps. As described in the next section, several probes made of different heat conducting materials were also actually designed (in particular, of wood and of silicon) and to comply with these experiments, those geometries were also analysed in the simulations. The results of the calculations on two suchlike probes are illustrated by Fig. 6; Fig. 7 gives a comparative summary of numerical, theoretical and experimental results. IV  EXPERIMENTS The particle velocity probe was placed in a 25 cm long standing wave tube [2,6] with a pressure microphone connected in the tube end, and at the other end a loudspeaker generating a broad band (0-10 kHz) signal. The probe measured the generated standing wave pattern and was rotated in steps of 3.75o. The measured frequency response was in correspondence with previous measurements on the sensitivity of the particle velocity sensor [1,9,4], showing a low-pass frequency response with a roll-off above 2 kHz. The measurements showed further that the polar pattern of the directional sensitivity was nearly frequency independent in the range from 50-1000 Hz for all investigated probes. To distinguish between the temperature induced and flow induced effects, two different series of probes were designed, one set with silicon as the surrounding material, the other with balsa wood. This material was chosen because of its extremely  low  heat  conductivity  compared  to that of  0 0.5 1 1.5 2 2.50102030405060gap distance (mm)angular shift ∆  Flow profile (theory)temperature (theory)both effects together (theory)flow profile (num.calc.)temperature (num.calc.)both effects together (num.calc.)experiment k=kSiexperiment k=kBa Figure 7. Angular shift as a function of gap distance (2lx) for different types of probes (with balsa wood, and with silicon as surrounding material), and compared with numerical calculations on the devices and with the model. The effects of the temperature, and of the disturbed flow profile, were separately investigated and then adde. silicon. Each set of probes comprised of devices with varying gap distances, with the full gap distance 2lx varying between 0.4 and 2.0 mm. Comprehensive comparison of these measurement results with both the numerical simulations and the theoretical calculations is made in Fig. 7. We see that the balsa based probes show significantly less temperature-related effects, resulting in a smaller angular shift. V CONCLUSIONS The sensitivity of a novel micromachined acoustic sensor consisting of four heated wires was analyzed theoretically and experimentally. The presented model is seen to be appropriate to describe the angular dependence and the magnitude of the sensor response and in satisfying correspondence with simulations adexperiments. Besides, the observed angular dependence of the sensitivity is frequency independent and cantherefore easily be corrected for. It is also concluded that the perturbed air flow due to the chip surface is the dominant reason for the observed angular sensitivity. REFERENCES [1] H.E. de Bree, P.J. Leussink, M.T. Korthorst, H.V. Jansen, T.S.J. Lammerink, and M. Elwenspoek, The Microflown: a novel device measuring acoustical flows, Sensors and Actuators A 54,  552-57, 1996 [2] H.E. de Bree, The Microflown Ph-D Thesis, University of Twente, the Netherlands (90-36509262), 1997 [3] H-E. de Bree, T.S.J. Lammerink, M.C. Elwenspoek, J. Fluitman, Use of a fluid flow-measuring device as a microphone, Patent PCT/NL95/00220, 1995  [4] D.R. Yntema, J.W. van Honschoten, H.E. de Bree, R.J. Wiegerink, M. Elwenspoek, A three dimensional Microflown, Proc. of MEMS Conf.(Istanbul) 654-7, 2006 [5] D.R. Yntema, J.W. van Honschoten, R.J. Wiegerink, Integrated 3D Sound Intensity sensor with four-wire particle velocity sensors, Proc. of DTIP, France, 168-170, 2008. [6] W.F. Druyvesteyn, H-E. de Bree, A new sound intensity probe; comparison to the Bruel & Kjaer p-p probe, J. audio Eng. Soc., 48, nr 1/2, 2000. [7] V.B. Svetovoy, I.A. Winter, Model of the Microflown microphone, Sensors and Actuators A 86 171-181, 2000 [8] J.W. van Honschoten, D.R. Yntema, V.B. Svetovoy, M. Dijkstra, R.J. Wiegerink, Analysis of the performance of a particle velocity sensor between two cylindrical obstructions, J. of Ac. Soc. of Am., 121, 2711-22, 2007 [9] J.W. van Honschoten, V.B. Svetovoy, G.J.M. Krijnen, M. Elwenspoek, Analytic model of a two-wire thermal sensor for flow and sound measurements, J. of Micromech. Microeng.,14, 1468-77, 2004 [10]  L.D. Landau, Course of Theoretical Physics, 6 Fluid Mechanics, 2nded., Butterworth Heinemann, Oxford, 2003 [11] H.Lamb, Hydrodynamics, 6thed. Cambridge Press, 1975. 

Modeling and characterization of the sensitivity of a hot-wire particle velocity sensor

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A new sound intensity probe; comparison to the Bruel & Kjaer p-p probe,

A three dimensional Microflown,

Analysis of the performance of a particle velocity sensor between two cylindrical obstructions,

Modeling and characterization of the sensitivity of a hot-wire particle velocity sensor

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