How to Take Read With Anemometer Intake

MEMS force per unit area and menstruum sensors for automotive engine management and aerospace applications

D. Sparks , in MEMS for Automotive and Aerospace Applications, 2013

4.four.2 Hot wire flow sensors

Thermal anemometers, or hot wire flow sensors, measure out the flow rate by monitoring the amount of estrus removed from a surface using one or more simple temperature sensors. They take found broad apply in the automotive industry for monitoring the air intake of the engine.

I of the simplest hot wire sensors is the Pirani approximate or thermal electrical conductivity device. The Pirani gauge is made up of a single wire or metal strip. The resistance of the wire is a function of the pressure level of a specific, stagnant gas surrounding the wire. This device is used to measure pressure or gas concentration in binary mixes, not catamenia, and has constitute widespread use with vacuum systems as a pressure gauge, but has not been unremarkably used in automotive and aerospace applications.

The hot wire menses sensor (Huijsing et al., 1982; Lambert and Harrington, 1986; Foss et al., 2005; Shmid et al., 2008) operates on a similar principle, in which the surrounding fluid interacts with a hot wire of the thin metal film. The cooling effect of a single wire tin can too exist used to determine the catamenia rate of the fluid passing over the hot wire. A single wire cannot indicate the direction of the flow, but this was not an issue with the early hot-wire flow sensors used in automotive applications for air intake. The calibration coefficients of this type of sensor likewise depend on the type of fluid flowing over the hot wire. In industrial applications, dissimilar coefficients are used for different gases. For the air intake application, this is mostly not an issue, although large differences in humidity touch the accuracy of the sensor.

By employing multiple wires or thin films, the flow rate and direction tin be measured. This is useful in four-cylinder engines where opposite pulsations, also called backflow, occur under heavy load. The backflow frequency can be 100   Hz, requiring special flow sensors. Figure 4.seven illustrates a bi-directional thin-film mass air menstruum sensor designed for this fast pulsating automotive flow sensor application (Lambert and Harrington, 1986). The central wire or metal strip is heated, while the 2 outer wires are sparse pic temperature sensors. These thin film temperature sensors are just xx to fifty microns from the heated metallic strip. A thermally insulating substrate is used to ameliorate thermal response time. A bypass sensor design also improved the performance of the sensor. As is illustrated in Fig. 4.7, as air flows in one direction the downstream temperature sensor heats up more than than the upstream temperature sensing strip. This allows for bi-directional sensing capability.

4.7. A bidirectional thin-flick thermal anemometer nether no flow (top) and directional menstruum conditions (lesser).

Thin film hot sparse film menstruation sensors have also been practical to fuel injections systems (Shmid et al., 2008). Figure four.8 shows an case of a sparse film hybrid co-fired ceramic flow sensor congenital into a fuel injector. The resistance change in a Wheatstone bridge is monitored every bit an indicator of fuel flow rate. This has been used in common rail straight diesel fuel injectors operating at pressures as loftier equally 1350   bar. Flow rate response times of 0.three to 1.5   ms could be detected, providing straight data on the hydraulic status of the interior of the fuel injector nozzles.

4.viii. (a, b) Hybrid ceramic hot wire menstruum sensor integrated into a fuel injector (Shmid, 2008).

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Book ii

Sjoerd Haasl , Göran Stemme , in Comprehensive Microsystems, 2008

2.07.2.three.three Calorimetric flow sensors (alternative names: thermotransfer, thermal dilution)

This class of thermal flows sensors measures flows past using at to the lowest degree 2 elements: a heater and a downstream sensor. The sensor measures the temperature change of the fluid by the heater. The thermal capacity of the fluid is an important parameter, hence the name calorimetric. As mentioned above, thermal anemometers tin can use distinct heater and sensor elements, simply in that case the sensor chemical element measures the temperature of the same thermal entity, measuring the oestrus loss without the heat passing through the fluid to go to the sensor. Platonic calorimetric menses sensors accept heater and sensor in perfect thermal isolation from each other except for the fluid. In reality, however, a big grayness zone exists, and the conduction of rut from the heater to the sensor via the substrate produces an unwanted signal. Table 3 provides an overview of micromachined calorimetric flow sensors.

Nearly always a second sensor is placed upstream. This makes the flow sensor bidirectional and provides a differential betoken, and then that the effect of the conduction through the substrate can be removed. An first is all the same created in cases where the temperature behavior of the upstream and downstream sensors differs. For resistive sensors, for example, this results in a difference of resistance. By placing a pair of sensors orthogonally to the offset pair, a 2d menses sensor is obtained (Fürjes et al. 2004, Kim et al. 2004, Mayer et al. 1996, Park et al. 2003a, Robadey et al. 1995).

A subclass of calorimetric flow sensors uses multiple hot sensors: heaters that double as sensors measure the deviation in estrus loss between the upstream and the downstream heater (Lammerink et al. 2000, Lyons et al. 1998, Nagata et al. 2001, van Oudheusden et al. 1989, Stenberg et al. 1988). What differentiates these sensors from the full Wheatstone bridge sensors described above is that the difference in temperature betwixt the upstream and the downstream heater is a issue of the catamenia-induced heat transfer between them, one heating the other through the fluid, and not from the difference in hydrodynamic boundary layer thickness. With this method, too period velocity, other parameters, such as, for example, in the case of Stenberg et al. (1988), the thickness of fouling biofilms growing on the sensor surface, can exist measured.

2.07.2.3.3.(i) Calorimetric period sensors: modes of operation

As with thermal anemometers, the heater can be run in different operating modes.CC, CV, and CP are like shooting fish in a barrel to implement but accept disadvantages. Lammerink et al. (1993) calculated the behavior of a three-wire calorimetric flow sensor in a microchannel (cross-channel span, see Figure 4(i) ), operated in CP mode. The temperature is assumed to exist abiding along the wire and to decrease linearly in the direction perpendicular to the menstruation and the wire. This results in the addition of a term $-\kappa T/l_z^{\mathrm{two}}$ to the steady-state ( $(\partial T/\partial t=0)$ ) 1D estrus transfer equation forth the length axis of the channel:

$U\frac{\partial T}{\partial x}=\kappa \frac{{\text{∂}}^{2}T}{\text{∂}{x}^2}-\kappa \frac{T}{l_z^2}$

with l_z being the height of the aqueduct. With purlieus conditions set to a heating ability P for the heater width 2L and ${\text{lim}}_{x\to \text{∞}}T=0$ , the solution of this equation becomes T  = T ₀ for −L<x<L, $T=T_0\exp \left({\text{λ}}_1\left(\mathrm{ten}+L\right)\right)$ for 10  > L (downstream), and $T=T_0\text{exp}\left({\text{λ}}_{\mathrm{two}}\left(x-L\right)\right)$ for 10  < L (upstream), with

${\text{λ}}_{\mathrm{ane},\mathrm{ii}}=\frac{\mathrm{ane}}{2\kappa }\left(U\pm \sqrt{U^2+4{\kappa }^{\mathrm{ii}}/l_z^2}\right)$

$T_0=\text{<mi mathvariant="italic" is="true">Pf</mi>}\left(k,{\text{λ}}_{1,2},\text{geometry}\right)$ , where f is a function of the geometry (height and width of the aqueduct and width of the heater), the thermal conductivity, k, and λ_1,2. In Figure 11 , the results are plotted for some typical values of U.

Figure eleven. Temperature profile near the heater for iii dissimilar menstruum rates. The vertical line indicates the position of the downstream sensor.

Figure 12 illustrates how, at low flow velocities, the temperature difference between the downstream and the upstream sensors, ΔT, increases linearly. However, at increasing velocities, the heater is cooled so much that ΔT decreases again.

Figure 12. Sensor temperatures every bit a function of the catamenia velocity U. T _h is the heater temperature, T _u and T _d the upstream and downstream sensor temperatures, respectively, and ΔT the temperature difference between the downstream and the upstream sensor. (Source: Lammerink T, Tas N, Elwenspoek M, Fluitman J 1993 Micro-liquid flow sensor. Sens. Actuators A (Phys.) A37–A38, 45–50. © 1993 Elsevier.)

CT control of the heater overcomes this problem and provides a monotonously increasing ΔT. Equally with thermal anemometers, it is also possible to control the temperature of the heater to maintain a CTD with the fluid.

The same TBA operating mode (Lammerink et al. 2000, Makinwa and Huijsing 2002a) is based on two heated elements at a small altitude from each other kept at the same temperature by a CP controller. The total ability to the two heated elements remains constant and the ratio of the power distribution betwixt the ii elements is a mensurate for the flow. Because of this, the restraint on the response of the temperature sensors is loosened, so that the two sensors only have to have the same response, unlike CTD where the response of the two sensors has to be the same at both the ambient and the heated temperature. This gives the freedom to use sensing methods that are highly sensitive and reproducible but nonlinear, such equally metallic–semiconductor thermocouples. An Ac-driven variant was presented by Lammerink et al. (2001) that consisted of two heating elements driven at 90° phase shift (creating a pair of 180° stage-shifted thermal signals) and a third element between them to sense the temperature variations using a phase-lock amplifier. The advantages of this arrangement lie in the avoidance of DC commencement and one/f noise.

Although nigh often the heater and the sensor blazon exercise not direct crave a specific operating mode, a specific bracket of sensors based on the pyroelectric issue does. Pyroelectric materials form a bracket of piezoelectric materials that distinguish themselves by generating a surface charge when undergoing temperature changes. Figure 13 shows a sketch of the device invented past Zemel and Rahnamai (1980) in which the electrodes on a lithium tantalate (LiTaO₃) substrate enable readout of the local charges upstream and downstream of a heating resistor. Past applying a constant-amplitude low-frequency voltage (i–three   Hz) on the heating resistor, a signal at the double frequency is generated at both electrodes. The downstream electrode will feel a larger temperature fluctuation and and so generate a larger voltage fluctuation. Advantages of the method are its low power consumption and large measurement range (0.5   ml   min⁻¹–20   fifty   min⁻ⁱ was reported in Hsieh et al. (1991)). The voltage is measured differentially over the two capacitors. A sparse-movie pyroelectric flow sensor was presented by Polla et al. (1983). Sputter-deposited ZnO was used, and the charge buildup was measured in DC mode. The charge decay was plant to be 85   h. A disadvantage of both variants, however, is the large time abiding, which ranges from seconds to tens of seconds. A thorough overview of pyroelectric flowmeters is given in Hsieh and Zemel (1995), Hsieh et al. (1995a, b).

Effigy xiii. Menstruum sensor based on the pyroelectric consequence. (Source: Zemel and Rahnamai H 1980).

SAW sensors, which are described in Section 2.07.ii.3.ane, accept also been used in calorimetric mode. Brace et al. (1989) used two SAW chips, thermally and acoustically isolated from each other, operating the upstream oscillator at a loftier CP and the downstream oscillator at low ability.

Recently, a new application surface area for calorimetric flow sensors has emerged. Instead of avoiding natural convection effects, the heater is heated up to a high temperature (150–700°   C) to create fluid menses abroad from the heater. If placed in an enclosed cavity, the management of the flow is determined only by the dispatch strength acting on it, and the airflow becomes a measure out for the acceleration and the inclination. The principle was first patented by Plöchinger (1994), and micromachined variants followed soon later (Billat et al. 2001, Chaehoi et al. 2005, van Honschoten et al. 2000, Leung et al. 1997, Lin and Jones 2005, Luo et al. 2001, Milanovic et al. 1998). With the exception of the microflown-based sensor described by van Honschoten et al. (2000), all are based on the three-element caloric principle: i heater in the center and one temperature sensor on either side in a bridge formation over a cavity ( Figure xiv ) The footstep to two-centrality designs is not big and has been demonstrated both for inclinometers (Billat et al. 2001) and accelerometers (Leung et al. 1998). Luo et al. (2002) accept presented an elaborate analysis of the parameters affecting the linearity, sensitivity, and frequency response of convective accelerometers.

Figure fourteen. Thermal accelerometer. (Source: Leung et al. 1998).

ii.07.2.3.3.(ii) The microflown

The microflown, conceptualized past de Bree et al. (1995), is a menstruum sensor designed to mensurate particle velocity. This quantity, defined as the flow volume rate per unit surface area, is used in acoustic theory together with acoustic pressure to draw acoustic waves. Since normal microphones measure acoustic force per unit area, the combination with a microflown allows the measurement of audio-visual intensity, the product of both. Combining a pressure level microphone with a microflown provides a low-price alternative to the conventional method of measuring catamenia intensity using two phase-matched microphones, as the phase-matching circuitry demands are loftier and, hence, expensive (Elwenspoek 1999). Conversely, two microflowns can exist placed at a distance from each other and have their signals combined to calculate the acoustic pressure (de Bree et al. 1996). The operating principle can exist applied to different calorimetric set-ups: a heater with 1 or two sensors or a ii hot-sensor principle (de Bree et al. 1995, van Honschoten et al. 2005). In contrast to pressure level microphones, the particle-velocity measurement is management sensitive and does not take a low-frequency cutoff. de Bree et al. (1999) used the common-fashion and differential signals from a microflown operated at CP to create a large-range management-sensitive flow sensor. The magnitude of the common-manner signal is used as it shows the typical flow dependency of a CP-operated thermal anemometer ( $P/\left(A+B\sqrt{U}\right)$ ) but is insensitive to the menstruation direction. The differential-mode point behaves like a typical nonuniform calorimetric menses sensor (run into Effigy 12 ), in which the sign too equally the highly sensitive depression-menstruation signal can exist used to complement the common-manner signal.

Table iii. Overview of calorimetric flow sensors

Reference	Heater	Sensor	Insulation a	Fluid temp. Compensation	Fluid type b	Manner	Sensitivity c,d,e	Sensitivity range f	Measured range	Resolution	ΔT/I/P	Time one thousand
Ashauer et al. (1999)	Polysilicon resistor	Al/Si thermopiles	SiO/SiN membrane	Yes: resistor on bulk	Water, oil, IPA	CP	1.7Fivemm −1 south	0.1   –   2.5   mm   south−1 in a 0.4   ×   0.vi   mm2 flow channel	−10   –10   mm   southward−i	0.1   mm   south−i	5   mW
Ashauer et al. (2001)	Polysilicon resistor	Al/Si thermopiles	SiO/SiN membrane	Yep: resistor on bulk	Water, oil	Non specified	0.66   V   ml−1 h differential		±0.25   ml   h−1	ten   μl h−1		2   ms
Billat et al. (2002)	Silicon resistor	Silicon resistor	Hotwire configuration	No (external temp. must be kept constant!)	Air, SFsix	CP (free convection measurement)
Brace et al. (1989): Dual-oscillator	SAW	SAW	None, hot flake: LiNbO3 substrate	No	Nitrogen	CP	−2.1kHz (cms −one )−0.5	half-dozen  – 500sccm	0   –   500   sccm		47.5°   C	10   –   30   s
Bracio et al. (2000)	Pt resistor	Ge thermistor	Membrane on air (textile non specified)	Yeah: two Ge thermistors on bulk		CT	No measurements		0.1   –   fifty   g   s−1	1.3   ml   h−i	20   K	1   ms in 50   mm dia. channel
Bruschi et al. (2005)	Two polysilicon resistors	Al/Si thermopiles (twenty n+polysilicon/Al thermocouples)	SiO membranes	No	Nitrogen	ΔT   =   0 method	0.35   mV at 200   sccm over thermopiles		0   –   200   sccm		6   5
Buchner et al. (2006)	W90Ti10 resistor or polysilicon resistor	W90Tix/Polysilicon thermopile	SiN membrane on polymer filling	No	Water	CT	3V  μl −1 s	0   –   2   μl   south−1	0   –   twenty   μl   s−1		44.i   K
Dillner et al. (1997)	NiCrSi resistor	Bi0.87Sb0.xiii thermopile	SiN/SiO/SiN stress comp membrane	Yes – cold junction on bulk	Dry air	CP	5.nine   Five   West−i  g−i  s		0   –   1660   sccm		1   mW
Ernst et al. (2002)	Cr resistor	Ge thermistor (amorphous)	SiN membrane	Yes: 2 Ge thermistors on bulk	Water	CP (isocaloric) and CT	CT: 2.197   V2 (mg   s−i)−0.5	CT: 0   –   7   μl   due south−ane	100   nl   h−1  –   3   μl min−1		CT: five   Thousand
							CP: 1.07   K   W−1 μl−1 h	CP: 0.1   –   90   μl h−i			CP: five   mW
Frederick et al. (1985)	NiCr resistor	LiTaOiii substrate with 2 NiCr electrodes on either side and one on the lesser: pyroelectric	None (LiTaOiii wafer)	No (just stated to be "relatively contained of fluid temperature")	North2,Otwo	AC voltage	20   mV at 500   yard   min−1		0.v   –200   yard   min−1
Fürjes et al. (2004)	Pt resistor	Pt resistor	SiN membrane	No	Air	CC	2   mV at 4   g   south−1		0   –   4   m   s−1
Glaninger et al. (2000)	Pt resistor or NiCr resistor	Ge thermistor (amorphous)	SiN membrane	Yes: two Ge thermistors on bulk	Air	CTD (CP for label)	ΔP   =   24   mW at fifty   ml   s−1 in 0.54   mm 2 channel		0.01   –   200   1000   due south−1 in 50   mm diam tube		25   K	CT: 20ms
									0.6   ml   h−one  –   150   l   h−1 in 0.54   mm2 cross-sectional tube			CP: ane.6ms
Gongora-Rubio et al. (1999)	Ru-based resistor	NTC thermistor (b   ≈   2000)	None (Green Record is thermally conductive)	Yes: separate unheated bridge	Nitrogen	CP	ΔT=18   K at iv   slm for threescore   mA		0   –   iv   slm			1   s
Häberli et al. (1997)	Polysilicon resistor	north-polysilicon/p-polysilicon thermopile (20)	SiO membrane on air	Yeah: thermopile common cold on majority	Air	Not specified	38   mV at 38   m   due south−1		0   –   38   m   s−1
Johnson and Higashi (1987)	NieightyFe20 (permalloy) resistors	NilxxxFe20 (permalloy) resistors	SiN membrane on air	Yes: resistor on bulk (upstream)	Air, He, COtwo	CC	0.037   Five g−1 min air	0   –   1   g min−1	0   –   one.ii   one thousand min−1 (i.e. For air at STP = 1 fifty min−ane)			3   ms
Kaltsas and Nassiopoulou (1999)	p-polysilicon resistor	Al/p-polysilicon thermopile	forty   μm thick porous silicon layer	Yes: thermopile cold on bulk	Nitrogen	CC	half dozen   mV   m−1  s   W−1	0   –   0.four   thou   s−1		4.one   mm   due south−1 ∼ ten   sccm		1.5   ms
Kaltsas et al. (2002)	p-polysilicon resistor	Al/p-polysilicon thermopile	40   μm thick porous silicon layer	Yes: thermopile cold on bulk	N2, O2, Ar	CC and CP (evaluated)	Nii: 1.045   mV (m   s−1)−0.5		0   –   iv   m   s−1			1.5   ms
							O2: 1.105   mV (g   s−1)−0.5
							Ar: 0.766   mV (thousand   southward−1)−0.five
Kim et al. (2004)	Pt resistor	Pt resistor	Si membrane (30   μm)	No	Air	CT – only flow direction is calorimetric!	0.32   Five at 10   m   s−1		0   –   x   m   s−1	0.v   one thousand   s−i
Koch et al. (1999)	Ti resistor	Ti resistor	SiN/SiO grid in the liquid	No	Ethanol	No measure out-ments	No measure-ments
Lai et al. (1997)	Diffused resistor	Diffused resistor	None	Yes: divide chip on back side of PCB	Gas	CTD	one.37 at 3   m   s−1		−4 to 4   1000   south−1		44.5   K
Lammerink et al. (1993)	Cr/Au resistor	Cr/Au resistor	SiN bridge	No	Water, air, IPA	CV	xiii   μV   K−1, IPA: viii   ×   teniii1000   thou−one  s	IPA: 0   –   1.iv   mm   southward−1	0   –   ii   grand   southward−1		3   V
							h2o: 2   ×   10threeK   chiliad−1  s	h2o: 0   –   2   mm   s−1
							air: three   ×   teniiiG   m−i  s	air: 0   –   two   k   due south−ane
Makinwa et al. (2001)	Polysilicon resistors	Al/p+Si (diffused) thermopile	None (Bipolar process)	Yeah: off-flake diode	Air	CTD	Not specified		0.ane   –   25   m   s−1	0.5   thou   due south−1 (±3%), 3°	Normally 15   K
Makinwa and Huijsing (2002b)	Polysilicon resistors	Al/p+Si (diffused) thermopile	None	Yes: off-fleck diode in case of CTD	Air	CTD and CP (TBA)	Non specified		two   –   18   1000   s−one	±four%, ±2°
Mayer et al. (1995): Air-gap	Polysilicon resistor	Al/polysilicon thermopiles	SiO membrane on air	Aye: thermopile cold on bulk	Nitrogen	CV	0.25   –   1.2   V   k−one  s   W−1		0.i   –   4   one thousand   s−1
Mayer et al. (1995): Bridge	Polysilicon resistor	Al/polysilicon thermopiles	SiO membrane on air	Yes: thermopile cold on bulk	Nitrogen	CV	0.039   –   0.56   Vm−1  s   W−i		0.1   –   4   m   southward−1
Moser et al. (1992)	Polysilicon resistor	Polysilicon resistor	SiO/SiO/SiN bridge	No	Nitrogen	CP	Type A: two.5   mV   m−ane  s linear, 180   μV   M−ane at 23   m   s−1 (nonlinear) Type B: 2.5   mV   yard−1  s linear	Type A: 0   –   5   m   s−1, Blazon B: i   –   3   k   s−1	Type A:0   –   23   m   south−aneBlazon B:0   –   25   chiliad   s−1	25   mm   s−one	Type A: 10   mW	4.4ms
Moser and Baltes (1993); Häberli et al. (1997)	Polysilicon resistor	due north-polysilicon/p-polysilicon thermopile (15) – Al at junctions to avert p-northward diodes	SiO membrane on air	Aye: thermopile common cold on bulk	Nitrogen	CP (equal mag. serial resistor on bulk)	three   μm CMOS: 0.36   mV   sccm−1  mW−i	iii   μm CMOS: x   –   100   sccm, 1.2   μm CMOS: 0   –   81   sccm in 0.5   mm dia. aqueduct	0   –   g   sccm in 0.5   mm dia. channel, Häberli: 0   –   38   yard   s−ane in open up air (half dozen   –   eight   mm opening)			1.45   ms
							one.2   μm CMOS: 0.v   mV   mW−1  sccm−1, Häberli:35   mV at 37   m   s−1
Nagata et al. (2001)	Pt resistor	Pt resistor	SiN membrane	No	Gas l	CV	one.85   Five at iv.75   m   s−one		0   –   0.8   m   s−i		ane.5   V
Nguyen and Kiehnscherf (1995)	Polysilicon resistor	Polysilicon resistor	Si membrane (10   –   30   μm) on air	No	Water, Nitrogen	CT and CP	water, CT: 0.26V (ml min −1 )−0.5		water:0   –   210   ml min−1		CT: 30   K	1   –   4   ms
							CP: nonlinear and not monotonous		nitrogen:0   –   300   ml min−1
							N2 CT: 2.nine   Vat 300   ml min
Nguyen et al. (2000)	Polysilicon resistor	Al/polysilicon thermopiles	SiN/ZnO membrane on air	Yep: thermopile common cold on bulk (out of menstruum!)	Air	CP	No data
Norlin et al. (1998)	Polysilicon resistor	Al/polysilicon thermopiles	Quartz substrate	Aye: separate on-scrap temperature sensor	H2o	Not specified	80   mV (amplified) at 150   μl min−1		0   –   200   μl min−1 in 1   ×   0.five   mm cantankerous-section channel
Oda et al. (2003)	Pt resistor	Pt/p+ doped polysilicon thermopile	SiO membrane on air	Yes: resistor on bulk	Air	CT	0.viiimV (l h −1 )−0.5		0.6   l   h−ane  –   12000   fifty   h−i		3.86   V
Ohnstein et al. (1990) (electronics past Chavan et al. (2003)	Pt resistor	Pt resistor	SiN membrane on air	Yep: two resistors on bulk	Air	CT	lxx   mV at 9000   thou min−1		0.9   –   9000   1000–min−ane
Pagonis et al. (2004)	p-polysilicon resistor	Al/p-polysilicon thermopile	x   μm thick porous Si on 10   μm thick air cavity	Yes: thermopile cold on bulk	Nitrogen	CT	14   mV (m   s−1)−0.5 W−1		0   –   2000   sccm (0   –   9.43   k   south−1)
Park et al. (2003a)	Pt resistor	Pt resistor	Si membrane	No – but not required either	Air	Not specified	Measures just angle		5 and 10   thousand   due south−1			tendue south
Polla et al. (1983)	Polysilicon resistor or P-implanted resistor	ZnO pyroelectric layer	None	No	Air	Not specified	184   mV at two   grand   south−1 8.eightmVthousand −1 southward	2   m   s−1		SNR   ≈   200	at ten.4 circuit gain and 36°   C	1 at ΔT   =   3°   C, forty   s at ΔT   =   10°   C for v   =   2   m   s−1
Qiu et al. (1996)	Ni resistor	Ni resistors	SiN membrane on air	Yes: ii resistors on bulk	Air (lx%RH)	CTl	700   mV at ii.7   m   s−1		0   –   3   yard   s−1		55   Yard
Rahnamai and Zemel (1981)	NiCr resistor	LiTaOthree substrate with 2 NiCr electrodes on either side and one on the bottom: pyroelectric	None (LiTaO3 wafer)	No (but stated to be "relatively independent of fluid temperature")	Air	AC voltage	10   mV2  cm−1 min	0   –   120   cm   min−1	0   –   2000   cm   min−1		3   Five
Rasmussen et al. (2001): polysilicon sensors	Polysilicon resistor	Polysilicon resistor	SiO membrane on air	No	Liquid	Non specified			0.01   mm   s−i
Rasmussen et al. (2001): thermopile sensors	Polysilicon resistor	Metal(Al 50 )/polysilicon thermopiles	SiO membrane on air	No	Liquid	Not specified	two.2   V   m−1  south	0   –   10   mm   due south−1 in 100   ×   100   μm channel				3   ms (response time)
Robadey et al. (1995): double bridge	Polysilicon resistor	Al/n-polysilicon thermopiles	SiO membrane on air	Yes: thermopile cold on bulk	Nitrogen	CP	0.19   V   grand−1  s   W−1	0   –   240   sccm	−400 to 400   sccm in a 0.2   ×   eight mmii channel	0.52   mm   south−1
Robadey et al. (1995): membrane (front mounting – rear mounting not analyzed)	Polysilicon resistor	due north-polysilicon/p-polysilicon thermopile – Al at junctions to avoid p–due north diodes	SiO membrane on air	No	Nitrogen	CP	1.9   V   m−1  s   W−1	0   –   200   sccm	−200 to 200   sccm in a 0.two   ×   5   mmii channel	5.2   mm   s−1
Rodrigues and Furlan (2003)	Polysilicon resistor	Polysilicon resistor	SiN membrane	No	Liquid	No measure-ments	No measure out-ments
Sabate et al. (2004)	Ni resistor	Ni resistor	SiN/oxynitride membrane on air	No	Air	CT	different resistor pairs: xl   mV   slm−ane	0   –   0.1 slm			150°   C
							two.1   mV   slm−1	0.1   –   2.0 slm
							0.25   mV   slm−ane	2.0   –   8.0 slm in 7   mm2 cross-department channel
Shin and Besser (2006)	Al resistor	Al resistors	Pyrex substrate	No	Nitrogen	CC	0.12Ω sccm−1 and 0.50   sccm−1 (unlike resistor configurations)	0   –   18   sccm in 610   μm x 500   μm aqueduct		±6.vi% at 200   mW (reproducibility test)	200   mW (42   K)	55.three   ms, response time=70   ms
Stenberg et al. (1988)	Implanted resistors	Diode	Polyimide-filled grooves, epoxy on back	No	Water	CC		0.2   –   2   yard   southward−1		±ten% no movie, ±30% with movie
Steurer and Kohl (1998)	Ge thermistor (amorphous)	Ge thermistor (amorphous)	Hotwire configuration (SiN support)	Aye: resistors on majority	Nitrogen, helium, hydrogen	CTD	Not specified		10   –   500   sccm	0.5%		<10   ms
Svedin et al. (2003a)	Polysilicon resistor	Polysilicon resistor	PI-filled groove	No	Air	Wheatstone bridge (CC)	0.seven   mV   Five−1 at 5   fifty min−i		0   –   xx   l   min−1
Tan et al. (2006)	Cr/Au resistor	Cr/Au resistor	Polyimide substrate	No	Nitrogen	CC	ΔR/R=8.5% at 2.75   50   min−1		0   –   two.75   l   min−i in 3   mm dia. tube
Tanase et al. (2002)	Polysilicon resistor	Al/Si thermopiles	polysilicon membrane on SiO	Yes: thermopile common cold on bulk, and (planned) separate temp. sensor	Blood	Non specified	14   mV at 21   cm   southward−1		0   –   fourteen   cm   s−i		2   K
Trautweiler et al. (1996)	Diffused resistor h	Al/Si thermopile	15   μm Si membrane on glass	No	Air	CCl	13.two   mV at five   slm in half-dozen   ×   1   mm2 channel		−five to 5   slm		(176   mW/13   M)
van Baar et al. (2001)	Pt resistor	Pt resistor	SiN beam	Not required due to AC measure-ment, alt. upstream probe	Gases: COii, N2	Not specified	Not specified		0   –   5   g   s−one
Weiping et al. (2005)	NiCr resistor	Ni resistor	None: SiN/SiO on Si substrate	No	Water	CV	0.55°   C at 350   μl min−i in single 130   μm dia. capillary 0.6°   C at 700   μl min−i in double capillary		10   –   700   μl min−i		∼   60°   C at 18   V
Yang and Soeberg (1992)	Diode	Diode	Si membrane (3   μm) on Glass substrate	Yes: upstream reference diode	Water	CP	32 V ml−1 min (amplified)	0   –   0.2   ml min−1	0   –   ten   ml min−1			0.two   s at 0.2   ml min−1
Yoon and Wise (1992)	Cr/Au resistor	Polysilicon/Au thermopile	SiO/SiN sandwich on air	No	Gas	N/A	Only flow direction sensor is calorimetric, the flow rate sensor is single chemical element
Yu et al. (1993)	NiCr resistor	LiTaO3 substrate with two NiCr electrodes on either side and one on the bottom: pyroelectric	None (LiTaO3 wafer)	No (but stated to be "relatively contained of fluid temperature")	Air	Air-conditioning voltage	Reynolds number sensitivity: 8   ×   10−5 at f   =   1.55   Hz

a

Silicon dioxide and silicon nitride are abbreviated to SiO and SiN, respectively, irrespective of stoichiometry.

b

Italicized fluid blazon indicates that the fluid was not specified in the reference.

c

Italicized sensitivities signal that the figure has been estimated from plots.

d

In cases where the output betoken is not described by a sensitivity factor, only the output at the largest measured flow charge per unit is given.

e

In cases of different parameters for like devices (e.g., temperature divergence), the highest response was given.

f

Sensitivity range: the sensitivity indicated is but valid within this range, if absent, the sensitivity is valid for the entire measured range. Italicized values signal that the figure was estimated from plots.

g

Time constants differ often between heating and cooling. The smallest fourth dimension constant is given. In case a pct is given, the value specifies the rise fourth dimension needed to reach this percentage of the steady-country value. Italicized values indicate that the figure was estimated from plots.

h

Causeless, non specified in the arich.

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Introduction

Hui Liu , in Wind Forecasting in Railway Engineering, 2021

1.3.ane.1 Anemometers selection

The ordinarily used anemometers contain thermal anemometer, ultrasonic anemometer, cup anemometer, Light Detection and Ranging (LiDAR), Sonic Detection and Ranging (SoDAR) etc. Co-ordinate to Chinese standard (TJ/GW089-2013), the anemometer tin can measure current of air speed within 0–60  grand/s and air current direction within 0–360°. These anemometers accept diverse characteristics as follows:

(a)

The thermal anemometer tin measure the temperature change of the hot wire, and then indirectly measure the current of air speed. In that location are two different types of thermal anemometers, including constant temperature type and constant current type, the latter of which is more widely used. The thermal anemometer is characterized by high sensitivity, small size, and fast response.

(b)

The ultrasonic anemometer can measure the air current speed by Doppler principle. The ultrasonic audio wave can spread faster in the downwind. According to the spread time difference of the ultrasonic sound wave between the downwind and headwind, the wind speed tin can be calculated. The ultrasonic anemometer has no mechanical wear, and fast response speed.

(c)

The cup anemometer can observe single-signal air current speed by measuring the cup'south rotation. The measurement functioning is afflicted by cup shape, rotation radius, and so on [51]. Due to the inertia, the cup anemometer has slow response speed.

(d)

The LiDAR and SoDAR are two remote sensing devices for wind speed measurement. The LiDAR is based on the particle trajectory, while the SoDAR is based on the audio reflection [52]. These two devices tin reach three-dimension scan of the current of air field, and have loftier spatial and temporal resolutions.

In Red china's Lanzhou-Xinjiang railway, a redundant wind measurement arrangement is applied, which is composed of loving cup anemometers and propeller anemometers [53]. In German railway, ii ultrasonic anemometers are installed at each single measurement station [7]. This redundant design of the anemometers tin can improve system robustness. If ane of the anemometers fails and the wind speed data cannot be output, the other anemometer can be used to generate the wind speed measurement results. Besides, the difference between the output values of the ii anemometers can be used to determine whether the output of the anemometer is reliable. If the difference exceeds a certain limit, there is a problem with an anemometer. At this time, the condition monitoring method can be used to decide which anemometer has failed, and the output value of the other anemometer tin can be used every bit the air current speed measurement results.

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Experimental techniques

Yanqiu Huang , ... Zhixiang Cao , in Industrial Ventilation Design Guidebook (Second Edition), 2021

4.3.8.7 Calibration

Constant calibration of most velocity instruments is necessary. Hot-wire or other thermal anemometers crave frequent calibration due to their complex and sensitive nature. They usually accept a significant time alter in their metrological characteristics, a drift. Vane anemometers are not that sensitive but require checking at gear up intervals. The Pitot-static tube is the only exception. Due to its fundamental nature, it does not require calibrating. It also can be used as a reference for other meters. This applies to the tube itself, provided it maintains its original geometry unchanged. Notwithstanding, the manometer used in conjunction with the Pitot-static tube requires calibration.

A calibration facility must produce the desired velocity range for the meter to be calibrated. The air temperature should exist kept constant over the test to ensure constant density. For thermal anemometers and ultrasonic anemometers, velocity calibration only is non sufficient. They should besides be checked for temperature compensation. In the case of omnidirectional probes, sensitivity to catamenia direction should be tested. In the case of low-speed (thermal) anemometers, their self-convection fault should be measured, and, for instruments measuring flow fluctuation (turbulence), dynamic characteristics testing should be carried out as well (Melikov, 1997).

Manufacturers of thermal anemometers provide small rigs for their calibration. They typically consist of a nozzle, an air supply unit, and a regulating valve. The probe is placed into the nozzle jet. The reference velocity is calculated from the nozzle upstream force per unit area and nozzle characteristics. Due to its small size, this type of rig can exist used just for hot-wire or other thermal anemometers (Bruun, 1995).

A good method for a uncomplicated calibration facility is a arrangement where a constant airflow is produced by using two water containers and an arrangement of a near constant force per unit area head (Yue & Malmstrom, 1998). The constant water flow into the 2d container displaces an equal airflow out of the container (Fig. iv.28). With this system the difficult measurement of a small airflow is inverse into a much easier and authentic measurement of a small h2o flow.

Effigy 4.28. A uncomplicated calibration facility for minor thermal probes.

The calibration air flows through a thin tube. The probe is placed at the get out of the tube. When the tube is long enough and the tube catamenia is laminar, the reference velocity for calibration can be calculated from the theoretical, fully developed laminar velocity profile.

To calibrate larger sensors/instruments such as vane anemometers, a air current tunnel is required. A calibration wind tunnel consists of an open up or closed tunnel, a fan to deliver the air, a nozzle to shape the velocity profile, and a mesh arrangement to make uniform and reduce the period turbulence. It may exist necessary to control the air temperature in the tunnel past ways of a heating/cooling system. The musical instrument to be calibrated is placed downstream from the nozzle, where an nigh apartment velocity profile is formed. The reference velocity may exist based on the measured flow rate and a known nozzle velocity profile, or information technology may exist measured close to the calibrated sensor using a Pitot-static tube. High-quality calibration wind tunnels are expensive facilities requiring skilled personnel to carry out the calibrations.

Very low (below 0.v   chiliad   s⁻¹) velocities are difficult to achieve in normal wind tunnels. In this range a slide tunnel is a better alternative. In a slide tunnel the air stands nonetheless and the sensor moves with a constant velocity on a slide. The reference velocity is based on the calibration distance and the traveling fourth dimension. As the length and fourth dimension can be measured very accurately, the calculated reference velocity value is of expert quality. Convective currents inside the slide tunnel restrict the everyman calibration velocities to effectually 0.05   m   s⁻¹. On the other hand, high velocities are restricted by the length of a straight tunnel, because a certain length of the tunnel is required for dispatch and deceleration of the slide.

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A thermistor-based instrument for measuring vehicle cooling airflow

D. Burton , J. Sheridan , in Vehicle Thermal Management Systems Conference and Exhibition (VTMS10), 2011

2 THERMISTOR ANEMOMETRY

Thermistors can provide velocity measurements in response to changes in heat transfer, which for a given geometry is principally a part of the air speed and temperature. The advantages of thermistors over other thermal anemometers such as hot-wires include accuracy at depression velocities, robustness and stability. The small size and low cost of thermistors ways that, despite being an intrusive device, a large number tin can be used across the radiator airplane. The high device resolution that tin can be obtained with thermistor anemometry is seen as a major advantage of this technique.

Thermistors are semi-conductors that showroom significant and precise changes of electric resistance in response to changes of their torso temperature. The thermistors that were used during this research were negative temperature coefficient (NTC) glass encapsulated bead thermistors.

Rasmussen [3] has shown that NTC thermistors showroom the following relationship between temperature and resistance (R_T ):

$R_T=R_0{e}^{\left[\beta \left(\raisebox{1ex}{$\mathrm{ane}$}\!\left/ \!\raisebox{-1ex}{$T$}\right.-\raisebox{1ex}{$\mathrm{one}$}\!\left/ \!\raisebox{-1ex}{$T_0$}\right.\right)\right]}$

Where R₀ is the thermistor reference resistance taken at a reference temperature T₀, and β is a textile constant. The temperature coefficient of resistance, α is:

[three] $\alpha =\left(\raisebox{1ex}{$1$}\!\left/ \!\raisebox{-1ex}{$R$}\right.\right)\left(\frac{\mathrm{dR}}{\mathrm{dT}}\right)=-\frac{\beta }{T^{\mathrm{ii}}}$

Where power is supplied to a thermistor by application of an electric electric current I, the power dissipated is P, and the temperature of the surrounding fluid is T_e, then the heat transfer equation becomes:

[three] $c\frac{\mathrm{dT}}{\mathrm{dt}}=P-\kappa \left(T-T_{\mathrm{east}}\right)$

Where c is the heat capacity, which is a property of the cloth and construction of the thermistor, and κ is the dissipation factor.

The dissipation cistron tin can be interpreted as the ability required to raise the temperature of a thermistor one degree to a higher place that of the surrounding fluid. The temperature of a thermistor, and therefore its resistance, will reply to a change in dissipation factor. It is this characteristic that makes thermistors appropriate for applications such equally anemometry. Where a thermistor is used in unidirectional laminar flow, with other fluid properties abiding, the dissipation factor can be regarded as a function of fluid speed only. Following Rasmussen [3]:

$T=T_{\mathrm{due\; east}}+\left[\raisebox{1ex}{$P$}\!\left/ \!\raisebox{-1ex}{$\kappa (U)$}\right.\right]$

Where thermistor reference resistance (R_east ) is taken at the fluid temperature, the thermistor trunk temperature tin be expressed:

$T=\frac{{\alpha }_e{T}_{\mathrm{due\; east}}^2}{\left({\alpha }_e{T}_e-\ln \left(\raisebox{1ex}{$R$}\!\left/ \!\raisebox{-1ex}{$R_e$}\right.\right)\right)}$

Equating the above ii equations yields:

$\kappa (U)=\left(\frac{P}{T_e}\right)\left[\frac{{\alpha }_{\mathrm{eastward}}{T}_{\mathrm{eastward}}}{\ln \left(\raisebox{1ex}{$R$}\!\left/ \!\raisebox{-1ex}{$R_{\mathrm{eastward}}$}\right.\right)}-\mathrm{i}\right]$

This equation provides a method for calculating the value of the dissipation factor in terms of simply determined values. Unfortunately, it is difficult to detect analytical solutions for the human relationship between fluid speed and dissipation factor as the shape, material-ratios and therefore thermal characteristics differ significantly betwixt individual thermistors. Measurements of the dissipation abiding, or the human relationship between thermistor resistance and flow velocity are all-time achieved empirically through calibration.

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Principles of air and contaminant movement within and effectually buildings

Alexander Zhivov , ... Yang Yang , in Industrial Ventilation Design Guidebook (Second Edition), 2020

7.iv.7 Applications of air jets

In industrial ventilation, both the experiment and CFD simulation accept been used for the written report of air jets. Conventional experiments used single-indicate measurements, include fast-response aerodynamic probe, ^225,226 thermal anemometer, ²²⁷ laser Doppler velocimetry (or laser Doppler anemonetry), ^228,229 etc. With advances in light amplification by stimulated emission of radiation imaging techniques, whole-field measurements have been developed recently, include particle-image velocimetry, ²³⁰ light amplification by stimulated emission of radiation-induced fluorescence (LIF). ^231,232 For the whole-field measurements, high-speed digital cameras were often used for recording the flows, to obtain the high resolution of the images.

Likewise the experiment, CFD simulation is also a competitive ways to narrate the free jet flows for its convenient and rapidity. RANS turbulence models, ^233,234 LES, ^235,236 and direct numerical simulation (DNS) ^237,238 were often used to solve the turbulent free jets. For RANS models, only time-averaged properties of the menstruum are required to be known, which may discard the details of the flow in the instantaneous fluctuations. While in DNS, no turbulence models are required, and the whole spectrum of turbulent scales are direct resolved, which results in huge computing burden. To balance the accuracy and computing load between the RANS and DNS models, LES was adult, in which large eddies are resolved directly and small eddies are modeled. Even though, CFD simulations for jets still need a big amount of calculating burden to capture the feature of the menstruum.

Recently some pioneering studies started to use POD to reconstruct the air flow jets. Lam ²³¹ used POD to extract the crucial modes for jet reconstruction, and compared the velocity fields from POD with that from the LIF experiment. Results show that with sufficient POD modes, the reconstructed field can exist very similar with the experiments.

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Sensors and Auxiliary Devices

The reading text for this course was originally written by, ... Robert McDowall P. Eng. , in Fundamentals of HVAC Command Systems, 2008

Mass Flow Meters [xiv]

While mass flow meters are available for liquid period sensing (such every bit Coriolis force meters, athwart momentum meters), they are not commonly used in HVAC applications.

Having covered water catamenia meters let u.s.a. turn to the measurement of air flow.

Mass menstruation meters do exist for h2o systems only are not used in air systems the opposite is true for air systems. One type of mass flow meter is very popular in air menstruation measurement: the thermal or hot-wire anemometer.

The thermal anemometer (see Figure iv-34 ) uses a heated probe placed in the air stream that is cooled past the motility of air in direct proportion to its mass flow rate. The probe consists of a temperature sensor and electrical resistance-heating chemical element. The device measures the electrical electric current required to keep the element at a abiding temperature (around 200°F) and translates this into a velocity betoken (assuming a abiding air density) that can be read on a meter or used in a control system.

Figure 4-34. Thermal Anemometer

In mod thermal anemometers, the heating element and temperature sensor are often the same device, usually a self-heated thermistor. Frequently, another temperature sensor is installed upstream of the probe to measure entering air temperature, used in the electronics to determine air density to provide a more than authentic velocity signal. These anemometers are called temperature compensated. The temperature sensor usually can be used past the control arrangement for other purposes too, often obviating the need to add some other air temperature sensor.

Thermal anemometers take the advantage of beingness able to sense much lower air velocities than pitot tube sensors. Common commercial models maintain ±ii% to ±iii% accuracy down to below 500   fpm; beneath this, accuracy drops to about x–20   fpm. A low limit of almost 100   fpm can be sensed with an accuracy of ±twenty   fpm. By comparison, pitot sensors are seldom authentic below about 400   fpm (0.01 inch wg) or even higher, depending on the transmitter used. Moreover, the pitot sensor volition not have the broad range of the thermal anemometer, which can read velocities from well-nigh 150 to 5000   fpm with reasonable accurateness. Pitot sensors are limited in range by the transmitter, which will have a range of no more than than almost 3 or iv to 1 while maintaining at least 10% accuracy. Therefore, thermal anemometers are a better sensor for some VAV systems that might feel a wide air flow operating range, and for outdoor air intakes on economizer systems where a broad range is typical. A relatively new product is a thermal anemometer designed to exist mounted in the fan inlet, a preferred location as described above for pitot sensors.

Thermal anemometers can also exist used to measure differential pressure across a barrier, such equally a wall between two rooms. Special mountings are used with a very pocket-sized porthole through the wall to let air to laissez passer. This velocity of this air flow is measured and can exist translated into a pressure difference past ways of the Equation 4-3 . The accuracy of the anemometer in this application is more than accurate than a differential pressure transmitter when the force per unit area difference being measured is very modest, less than virtually 0.02 inch wg.

An instance of an air menstruation meter is shown in Figure 4-35 .

Figure four-35. Air Flow Meter

Pitot tube sensors (run across Figure 4-36 ) are commonly used to mensurate the speed of aircrafts. They are also used very commonly for measuring both air and water menstruum in HVAC applications. The differential pressure in Equations 4-i–iv-3 in this example is the difference between the total and static pressures, a quantity called the velocity pressure level. Equally can be seen in Effigy 4-36 , the inner tube senses total force per unit area of the fluid, which due to the static pressure level plus the strength exerted past the fluid'south velocity, called the velocity pressure level. The outer tube has openings in the sides, which are not impacted by the fluid flow and therefore sense only static pressure. The difference between the 2 is used in Equations four-1–iv-3 to determine fluid velocity. In an ideal pitot tube, the C coefficient is a constant regardless of geometry.

Figure four-36. Pitot Tube Sensor

For air flow at standard density, velocity may be calculated from the pitot tube velocity differential pressure as:

(Equation 4-3) $V=4005\sqrt{\text{Δ}P}$

where ΔP is measured in inches of water gauge (wg) and V is measured in feet per minute (fpm).

Figure 4-37 shows an air flow measuring station (FMS) ordinarily used in duct applications.

Effigy 4-37. Air Flow Measuring Station

In large ducts, the FMS is composed of an array of pitot sampling tubes the pressure signals of which are averaged. This bespeak is fed to a square root extractor, which is a transmitter that converts the differential pressure bespeak into a velocity point. (With digital control systems, this adding can be made in software with improved accuracy over the use of a foursquare root extractor.) The velocity measured in this mode is an approximate average of duct velocity, only it is not a precise average because pressure level and velocity are non proportional (the average of the square of velocity is non equal to the boilerplate of the velocity).

Where fan air flow rate measurement is required, the preferred location of the pitot sensor is in the inlet of the fan. Two arrangements are bachelor. The first, which can be used with almost any fan, has two bars with multiple velocity force per unit area and static pressure ports mounted on either side of the fan axis. The second, much less mutual arrangement has multiple pinhole pressure taps that are congenital into the fan inlet by the fan manufacturer. Differential pressure is measured from the outer point of the inlet to the near constricted point, much similar a Venturi meter.

Locating the air flow sensor in the fan inlet has many advantages compared to a duct mounted pitot array. Offset, air flow is more often than not stable in the inlet (except when inlet vanes are used, in which case this location is non recommended) and velocities are high, which increases accuracy considering the differential pressure indicate will exist loftier. Even where inlet vanes are used, a location in the mixed air plenum space could be institute. This location besides reduces costs because the sensor array is smaller than the array required in a duct. Perhaps the nearly important advantage of this location is that it obviates the need to provide long straight duct sections required for the duct mounted assortment. The space needed for these duct sections seldom seems to be available in mod HVAC applications where the operating space occupied past HVAC systems is heavily scrutinized by the owner and architect, reduced to its smallest expanse possible, and consciously minimized.

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Volume 2

Ulrich Bonne , in Comprehensive Microsystems, 2008

two.11.2.v Conclusions

The to a higher place results and demonstrations indicate the feasibility and cost-effectiveness of coupling fluid motion induced by simple, available, and affordable membrane actuators to thermal or force per unit area microsensors to provide advantageous ways to determine:

(1)

Authentic, submillisecond response times of different flow-sensor structures and at different flows. At flows of up to ∼   1   m   southward⁻¹, the MB sensor menses response time decreases to 0.eight   ms, but the value for the more rugged Microbrick drops to 0.3   ms, and lower however at higher flows. Conservatively, this means that these sensors tin can be operated in oscillatory flows at frequencies of up to ∼   200 and 530   Hz, respectively. Application: Defining the actuator frequency range for the measurements discussed below, and generic response time limitations of such microstructure menstruation sensors.

(2)

Affordable composition correction factors for thermal and other flow sensors. The laborious arroyo used previously for thermal anemometers, based on the measurement of thermal conductivity and specific heat, tin can at present be replaced by one simple flow measurement to make up one's mind the volumetric composition correction factor, C _V, with a insufficiently negligible effort for calibration and electronics, that achieves at least ±iii% accuracy. Awarding: Flow sensing of fluids like gaseous or liquid fuels, whose multicomponent composition can vary versus time, course, and source.

(3)

The limerick of binary and higher-order mixtures. The determination of the correction factor, C _V, adds one easy-to-mensurate parameter to the choice of parameters bachelor to characterize the composition of gaseous or liquid mixtures.

(4)

Viscosity and other physical properties of fluids. With the laminar flow induced past a simple oscillatory membrane (EM rather than PE earphone speaker) operated beneath 100   Hz, we demonstrated a compact, light weight, and affordable arroyo to viscosity measurement. The measurable Air conditioning pressure drop signal serves indeed as a linear indicator of viscosity co-ordinate to Poiseuille's law for gluey flow, which is density independent if operated at depression kinetic energy conditions and besides pressure contained to the extent that the measured viscosity is pressure independent (∼   0.i%   bar). 1 driver for this evolution was to correlate viscosity (past itself or in combination with other parameters) to oxygen demand or Wobbe number.

Ii other classes of thermal sensors, also demonstrated with MB sensor chips, stand for a result of efforts to minimize sensor drift by shifting sensor readout from being dependent on analog changes in some physical sensor belongings, to frequency or digital changes in such sensors. But here they shall only be mentioned in passing:

(1)

Those geared to induce electronic RC type of self-oscillations past leveraging the temperature-dependent changes in the thin-film resistance of the sensor and the capacitance of the circuit. Such oscillations then change the sensor's output frequency betoken equally a outcome of changes in menstruation, thermal electrical conductivity, or specific heat. Consistent with the thermal response time, frequencies of upwards to 1000   Hz were generated (Bonne and Kubisiak 2001b). The generation of a frequency output may exist appealing, although does not eliminate the dependence of the underlying analog change in sensor resistance.

(2)

Those based on measuring the variable time delay associated with the change in heat transfer between the sensing elements involved in the measurement of the above thermal properties. Such digitally measurable delay, if primarily dependent on changes in flow rather than in the analog resistances of the sensor elements, would bring u.s.a. closer to the platonic digital sensor (Bonne and Kubisiak 2000, Bonne et al. 2000b, 2001b). In this promissing approach, one should however check and verify that the influence of the sensor analog resistances may notwithstanding play a role in the overall sensor operation.

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Characteristics Of Heat Transfer Coefficient Derived From Application Of Oestrus-Mass Transfer Analogy On Sublimating Naphthalene Deejay Data

Stephen KW. Chang Ph.D. , Richard R. Gonzalez Ph.D. , in Transport Phenomena in Oestrus and Mass Transfer, 1992

3.2 Experiment

Circular naphthalene disks were attached to five locations on a stationary, unheated, life-size (body surface area 1.68 k²) copper manikin in an environmental sleeping room with precise temperature and air circulation control. Figure one shows the positions of the naphthalene disks on the upper arm, lower arm, thigh, shin, and chest of the manikin. The ambient temperature (T_a) in the environmental chamber was set to 30°C, with dew point temperature at 5°C. The thirty°C setting was chosen to facilitate the rate of naphthalene sublimation. The bedroom and manikin were allowed to equilibrate for 24 hours at thirty°C before beginning the written report. The life-size manikin simulates the human body contours and avoid the problem of perspiration. A manikin also eliminated any insensible rut loss component associated with human being subjects.

Effigy 1. Manikin schematic diagram showing naphthalene deejay locations.

The environmental chamber employed the conventional aspirated airflow. Honeycomb grids, at the airflow entrance, directly the catamenia and minimize turbulence. Since the naphthalene disks were place frontally on the manikin (see Figure 1 ), the disks faced the airflow ordinarily (i.e. perpendicularly). The regional temperature and air velocity were measured individually at the five deejay sites. Omnidirectional thermal anemometers (range 0 – 3 g/southward) and thermistor temperature probes were placed approximately 2 cm abroad from and 2 cm above each naphthalene disk. The probes were placed equally close equally possible to measure the exact status experienced by the disks, but besides they must cause as little disturbance as possible to the impinging airflow. The free-stream air velocity in the sleeping room was measured with a loving cup and a heated-wire anemometer positioned approximately ane meter in forepart, at the caput level of the manikin. The duration of each experiment was 55 minutes.

The frontal placements of the naphthalene disks were chosen to enable comparison of our results with Sogin's [1958] and Sparrow and Tien's [1977] studies where air stream likewise impinged upon the naphthalene ordinarily. The naphthalene deejay casting cassettes for the upper arm, lower arm, thigh, and shin sites were appropriately curved to adapt to surface curvatures of corresponding limbs on the manikin. These iv naphthalene disks therefore exposed elliptic rather than circular surfaces. Casting cassettes used for the chest were flat. The elliptic area of each cassette disk was measured, included in Tabular array 1, and properly quantified (within the $\stackrel{•}{\text{m}}$ term in Eqn.{five}).

Tabular array i. Physical Attributes of Stationary Manikin and Naphthalene Disks.

Naphthalene disk site	Regional air velocity grand/due south	Air temperature °C	Disk area cm2
Upper Arm	0.605 ± 0.016	30.0 ± 0.11	19.2
Lower Arm	0.712 ± 0.011	29.9 ± 0.12	19.3
Thigh	0.417 ± 0.014	xxx.0 ± 0.13	18.9
Shin	0.687 ± 0.023	30.0 ± 0.11	nineteen.iii
Chest	0.489 ± 0.013	30.0 ± 0.13	xviii.1

The casting cassettes were modified from aluminum camera lens covers. During casting, the open up end of the lens cover was sealed with a nonadhering plate. The underside of the lens embrace has two drilled holes (0.5cm dia.), ane for pouring the naphthalene, the other served equally a vent. Scintillation grade naphthalene (melting point ≈ 80°C) was melted and then poured into the casting cassettes. The liquified naphthalene quickly hardened at room temperature, subsequently which the nonadhering plate was removed to reveal a smooth naphthalene surface. The pour and vent holes were covered with surgical record and the disks were stored individually in closed containers. All naphthalene disks were allowed to equilibrate in the chamber at thirty°C for 24 hours before using. A naphthalene deejay, referred to in this study, comprised of the modified photographic camera lens comprehend and the naphthalene within. For the written report, the disks were kickoff mounted into vinyl retainers, which were in turn fastened to the manikin using Velcro straps. The retainer vinyl plates have a circular hole cut out in the center. The diameter of the cut hole was slightly smaller than the bore of the deejay and so that the disk fitted tightly in position.

The naphthalene disks were weighed immediately before and after an experiment on a precision balance sensitive to ±0.01 mg. This precision balance weighing method of determining naphthalene mass transfer has been shown to be accurate to within i%-two% of the more elaborate methods such as integration of local surface depth measurements [Saboya & Sparrow 1974].

Every bit an effort to decide the relationship between h_n and h_c, a mass transfer experiment like to that performed by Nishi and Gagge [1970b] was employed. A one.21m long lite-weight magnesium bar was rotated at its center by a stepping electrical motor at eight constant speeds ranging from 9 to 30 rpm, in a still-air chamber. One flat naphthalene disk was attached to each finish of the magnesium bar. As the bar rotated, the naphthalene disks ever faced ordinarily (head-on) into the air. The rotation resulted in translational speeds from 0.548 to 1.83 m/s for the disks. Since the bedroom had yet-air, the translational speeds were also exactly the air velocities that the naphthalene disks were facing. Duration of each experiment was the same as the manikin studies (55 minutes). The two naphthalene disks were weighed before and after the experiment on the same precision balance.

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THE Consequence OF WALL ROUGHNESS ON AN OPEN Aqueduct BOUNDARY LAYER

M.F. Tachie , ... R. Balachandar , in Engineering Turbulence Modelling and Experiments 5, 2002

INTRODUCTION

Wall roughness is prevalent to some degree in most industrial and environmental flows. Furthermore, information technology is more often than not recognised to accept a significant influence on the transport properties of boundary layers, and has therefore been the subject area of extensive experimental study. All the same, any attempt to implement numerical models for rough wall turbulent near-wall flows volition immediately demonstrate that compared to the case of smooth surfaces, predictive methodologies for rough surfaces are clearly deficient. For instance, the skin friction at the surface becomes difficult to generalise in ways which require only a minimal description of the surface geometry. More importantly, the turbulence structure in a boundary layer on a rough surface differs from its analogue on a smooth surface in specific ways.

The growing torso of experimental evidence suggests that the wall similarity hypothesis which states that the effects of surface roughness are confined to the extent of the roughness sublayer is incorrect, and that surface roughness can influence sure aspects of the turbulent structure fifty-fifty in the outer region of a purlieus layer. For instance, Mazouz et al. (1994) reported measurements in a channel menstruation over rough surfaces and observed that surface roughness reduces the level of turbulence. Krogstad and Antonia (1999) performed cross-wire measurements on shine and rough surfaces in a turbulent boundary layer. Their results showed that the streamwise and wall-normal components of the turbulence intensity likewise as the Reynolds shear stress are college for the crude surfaces than the corresponding smooth wall data. The rough wall turbulent boundary layer study reported by Ligrani and Moffat (1985) showed that surface roughness increases the streamwise turbulence intensity, but the wall-normal turbulence intensity and Reynolds shear stress are independent of wall roughness. The summary discussion above reveals that our present noesis of roughness effects on nigh-wall turbulent flows is contradictory. Most of the rough wall data available in the literature were obtained using conventional thermal anemometers which may have been affected by the high local turbulence levels shut to the wall. Furthermore, the friction velocity U _τ is often used to calibration both the mean and turbulence quantities, notwithstanding an accurate determination of the wall shear stress (or friction velocity) remains an important inquiry topic especially for menstruation over rough surfaces. In most of the earlier rough wall boundary layer studies, the classical Clauser plot technique and Hama'due south (1954) formulation were used to decide the friction velocity. Nevertheless, the values of U_τ reported in some studies (eastward.g. Mills and Hang, 1983; Bandyopadhyay, 1987; Krogstad et al., 1992) using Clauser or Hama's formulation deviate significantly from those obtained from either a momentum residuum or by extrapolating the Reynolds shear stress to the wall. Bradshaw (1978) and Mills and Hang (1983) attributed the disparity to the neglect of the role of the wake component of the velocity profiles.

The nowadays paper reports LDA measurements in the boundary layer generated in a low Reynolds number open channel menses. Specifically, we examine the furnishings of surface roughness on the mean velocity profile, the streamwise and wall-normal fluctuating velocity components, the Reynolds shear stress and the triple correlations. An insightful presentation of these results requires that the correct scaling laws be used to analyse the experimental data. At the present time, different scaling laws take been advocated based on theoretical approaches which incorporate different assumptions. Nosotros have considered both the classical arroyo post-obit Millikan (1938) which scales the outer region of the mean flow using the friction velocity, U_τ, besides every bit the more recent approach advocated by George and Castillo (1997), future denoted by [GC], who utilise the external velocity, U_east for a nix pressure gradient turbulent boundary layer. For example, the classical arroyo leads to the post-obit logarithmic contour for the overlap region:

[1] $U^{+}=\frac{1}{\kappa }ln{y}^{+}+C-\Delta U^{+}$

where U = U/U_τ y⁺ = yU_τ /five, κ and C are universal log police force constants (i.eastward. independent of Reynolds number) and ΔU⁺ is the roughness shift.

The theory proposed past [GC], on the other mitt, leads to a power of the class:

[two] $U^{+}=C_i{\left(y^{+}+a^{+}\right)}^{\gamma }$

In Eqn (2), the coefficient C_i and the exponent γare power police force constants which depend on the local Reynolds number δ⁺ = δU_τ /5, where δ is the purlieus layer thickness, and a⁺ (= −16) represents a shift in the origin for measuring y, associated with the growth of the mesolayer region (thirty ≤ y⁺ ≤ 300). George and Castillo (1997) also proposed the post-obit composite velocity contour which gives a reasonable description of the mean velocity in the viscous sublayer, the buffer region, and the overlap region:

[3] $\begin{array}{ll}{U}^{+}& =\left[y^{+}+c_4{y}^{+5}+c_5{y}^{+5}\right]exp\left(-d{y}^{+\mathrm{half\; dozen}}\right)+\\ & C_i{y}^{+}\left[1+\gamma a^{+}{y}^{+-1}+\frac{1}{2}\gamma \left(\gamma -\mathrm{ane}\right)a^{+\mathrm{two}}{y}^{+-2}\right]\text{x}\left(1-exp\left(-d{y}^{+\mathrm{half-dozen}}\right)\right)\end{array}$

where d is a damping parameter (chosen as d = 8×x^−eight) to gear up the transition from the viscous wall region to the overlap region at y⁺ = 15. The LDA measurements reported by Eriksson et al. (1998) suggests that c₄ = −0.0003±0.0001 while [GC] recommended c₅ = 13.5×10^{−half dozen}. The theory proposed by [GC] also showed that the fluctuating velocity components scale on the external velocity U_e while the triple correlations scale on U_τ ²Ue. Although of import distinctions exist between a boundary layer in open channel catamenia and a approved null-pressure gradient boundary layer, e.thousand. the free surface upshot, at that place are stiff similarities which allow us to draw general conclusions regarding the furnishings of surface roughness on near-wall turbulence structure.

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