aerial infrared thermography abstract.pdf

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Optics and Lasers in Engineering 44 (2006) 261–281 Infrared thermography: An optical method in heat transfer and fluid flow visualisation T. Astarita , G. Cardone, G.M. Carlomagno Universita` degli studi di Napoli ‘‘Federico II’’, Dipartimento di Energetica Termofluidodinamica Applicata e Condizionamenti Ambientali, DETEC, P.le Tecchio 80, 80125 Napoli, Italy Available online 23 May 2005 Abstract This paper deals with the evolution of infrared thermography into a powerful optical method to measure wall convective heat fluxes as well as to investigate the surface flow field behaviour over complex geometries. The most common heat-flux sensors, which are normally used for the measurements of convective heat transfer coefficients, are critically reviewed. Since the infrared scanning radiometer leads to the detection of numerous surface temperatures, its use allows taking into account the effects due to tangential conduction along the sensor; different operating methods together with their implementations are discussed. Finally, the capability of infrared thermography to deal with three complex fluid flow configurations is analysed. r 2005 Elsevier Ltd. All rights reserved. Keywords: Heat-flux sensors; Convective heat transfer; Surface flow visualisation; Infrared thermography 1. Introduction Usually, measuring convective heat fluxes requires both a sensor (with its corresponding thermal model) and some temperature measurements. In the ordinary techniques [1–6], where temperature is measured by thermocouples, resistance ARTICLE IN PRESS 0143-8166/$ - see front matter r 2005 Elsevier Ltd. All rights reserved. doi:10.1016/j.optlaseng.2005.04.006 Corresponding author. Tel.: +39 081 768 3389; fax: +39 081 239 0364. E-mail address: [email protected] (T. Astarita).

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ARTICLE IN PRESS

Optics and Lasers in Engineering 44 (2006) 261–281

0143-8166/$ -

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Infrared thermography: An optical method inheat transfer and fluid flow visualisation

T. Astarita�, G. Cardone, G.M. Carlomagno

Universita degli studi di Napoli ‘‘Federico II’’, Dipartimento di Energetica Termofluidodinamica Applicata

e Condizionamenti Ambientali, DETEC, P.le Tecchio 80, 80125 Napoli, Italy

Available online 23 May 2005

Abstract

This paper deals with the evolution of infrared thermography into a powerful optical

method to measure wall convective heat fluxes as well as to investigate the surface flow field

behaviour over complex geometries. The most common heat-flux sensors, which are normally

used for the measurements of convective heat transfer coefficients, are critically reviewed.

Since the infrared scanning radiometer leads to the detection of numerous surface

temperatures, its use allows taking into account the effects due to tangential conduction

along the sensor; different operating methods together with their implementations are

discussed. Finally, the capability of infrared thermography to deal with three complex fluid

flow configurations is analysed.

r 2005 Elsevier Ltd. All rights reserved.

Keywords: Heat-flux sensors; Convective heat transfer; Surface flow visualisation; Infrared thermography

1. Introduction

Usually, measuring convective heat fluxes requires both a sensor (with itscorresponding thermal model) and some temperature measurements. In the ordinarytechniques [1–6], where temperature is measured by thermocouples, resistance

see front matter r 2005 Elsevier Ltd. All rights reserved.

.optlaseng.2005.04.006

nding author. Tel.: +39081 768 3389; fax: +39081 239 0364.

dress: [email protected] (T. Astarita).

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T. Astarita et al. / Optics and Lasers in Engineering 44 (2006) 261–281262

temperature detectors or pyrometers, each transducer yields either the heat flux at asingle point, or the space-averaged one; hence, in terms of spatial resolution, thesensor itself can be considered as zero-dimensional. This constraint makesexperimental measurements particularly troublesome whenever temperature, and/or heat flux, fields exhibit spatial variations.

As long as the fluid is transparent to the employed infrared band, the infraredscanning radiometer (IRSR) constitutes a true two-dimensional temperaturetransducer since it allows the performance of accurate measurement of surfacetemperature maps even in the presence of relatively high spatial temperaturevariations. Correspondingly, the heat-flux sensor may become two-dimensional aswell. In particular, infrared thermography can be fruitfully employed to measureconvective heat fluxes, in both steady and transient techniques [7–9]. Within thiscontext, IRSR can be intrinsically considered as a thin-film sensor [5] because itgenerally measures skin temperatures. The thermal map obtained by means ofcurrently available computerised thermographic systems is formed through a largeamount of pixels (20–300K and more) so that IRSR can be practically regarded as atwo-dimensional array of thin films. However, unlike standard thin films, which havea response time of the order of microseconds, the typical response time of IRSR is ofthe order of 10�1–10�3 s.

The use of IRSR as a temperature transducer in convective heat transfermeasurement appears, from several points of view, advantageous if compared tostandard transducers. In fact, as already mentioned, IRSR is fully two-dimensional;it permits the evaluation of errors due to tangential conduction and radiation, and itis non-intrusive. For example, the last characteristic allows to get rid of theconduction errors through the thermocouple or resistance temperature detectorwires.

2. Heat-flux sensors

Heat-flux sensors generally consist of plane slabs with a known thermal behaviour,whose temperature is to be measured at fixed points [1–6]. The equation for heatconduction in solids applied to the proper sensor model yields the relationship bywhich measured temperature is correlated to the heat transfer rate.

The most commonly used heat-flux sensors are the so-called one-dimensional ones,where the heat flux to be measured is assumed to be normal to the sensing elementsurface, i.e. the temperature gradient components that are parallel to the slab planeare neglected. In practice, the slab surfaces can also be curved, but their curvaturecan be ignored if the layer affected by the input heat flux is relatively small ascompared to the local radius of curvature of the slab.

In the following, first ideal one-dimensional sensors are considered and then,whenever possible, the use of some of them will be extended to the multi-dimensionalcase. The term ideal means that thermophysical properties of the sensor material areassumed to be independent of temperature and that the influence of the temperaturesensing element is not considered.

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The most commonly used one-dimensional sensor models are:

(1)

Thin-film sensor: A very thin resistance thermometer (film) classically measuresthe surface temperature of a thermally thicker slab to which is bonded. Heat fluxis inferred from the theory of heat conduction in a semi-infinite wall. The surfacefilm must be very thin so as to have negligible heat capacity and thermalresistance as compared to the slab ones. To use this sensor with infraredthermography, the heat exchanging surface must be necessarily viewed by IRSR.

(2)

Thick-film sensor: The slab is used as a calorimeter; heat flux is obtained from thetime rate of change of the mean slab temperature. This temperature is usuallymeasured by using the slab as a resistance thermometer.

(3)

Wall calorimeter or thin-skin sensor: The slab is made thermally thin (so that itstemperature can be assumed to be constant across its thickness) and, as in thecase of the thick-film sensor, is used as a calorimeter. Heat flux is typicallyinferred from the time rate of change of the slab temperature which is usuallymeasured by a thermocouple. To use this sensor with infrared thermography,either one of the slab surfaces can be generally viewed by IRSR.

(4)

Gradient sensor: In this sensor the temperature difference across the slabthickness is measured. By considering a steady-state heat transfer process, heatflux is computed by means of the temperature gradient across the slab. Thetemperature difference is usually measured by thermopiles made of very thin-ribbon thermocouples, or by two thin-film resistance thermometers.

(5)

Heated-thin-foil sensor: This method consists of steadily heating a thermally thinmetallic foil, or a printed circuit board, by Joule effect and by measuring the heattransfer coefficient from an overall energy balance. Also, in this case, due to thethinness of the foil, either one of the slab surfaces can be viewed by IRSR.

Strictly speaking, there is another type of one-dimensional sensor, the circularGardon gauge, in which the heat flux normal to the sensor surface is related to aradial temperature difference, in the direction parallel to the gauge plane [1]. Thissensor is practically of no use in infrared thermography.

Recently, another type of heat-flux sensor based on a three-dimensional unsteadyinverse model and IRSR surface temperature measurements has been also developed[10] but for sake of simplicity it will not be herein described.

Application of IRSR to both the thick-film and the gradient sensors is not verypractical, so these sensors will not be herein described. The heated-thin-foil sensorrepresents a quasi-steady technique that will be discussed in the next paragraph; thethin-film and the wall calorimeter sensors constitute transient techniques that will betreated in the following one.

3. The heated-thin-foil steady-state technique

Within the class of steady-state techniques to measure convective heat fluxesbetween a fluid stream and a surface, a method, where the application of IRSR seems

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Fig. 1. Heated-thin-foil sensor.

T. Astarita et al. / Optics and Lasers in Engineering 44 (2006) 261–281264

to be very effective, is the heated-thin-foil technique. The sensor is made of a thinmetallic foil which is heated by Joule effect (see the sketch of Fig. 1a). The mainlimitation of this technique is that, for practical reasons, the exchanging surfaceshould have a cylindrical, or conical, geometry.

In the following, it is initially supposed that the sensor is one-dimensional and thatthe surface not exposed to flow is adiabatic. By making a very simple (one-dimensional) steady-state energy balance, it is found

Qj ¼ Qr þQc, (1)

where Qj is the imposed constant Joule heating per unit area, Qr is the radiative heatflux to ambient, and Qc is the convective heat flux to fluid.

The radiative heat flux can be evaluated by

Qr ¼ s�ðT4w � T4

ambÞ, (2)

where s is the Stefan–Boltzmann constant, � is the total emissivity coefficient, and Tw

and Tamb are the temperature of the wall and of the experimental ambient,respectively. When standard techniques are used to measure the wall temperature, itis possible to have a very low wall emissivity coefficient so as to ignore the radiative

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heat flux to ambient. Obviously, this is not the case when measuring temperatures bymeans of IRSR.

The convective heat flux can be expressed according to Newton law:

Qc ¼ hðTw � T rÞ, (3)

where h is the convective heat transfer coefficient and Tr is a reference temperature.The reference temperature depends on the stream experimental conditions. Forexample, for high Mach number flows (or for the mixing of two streams at differenttemperatures), the correct choice is the adiabatic wall temperature [11–14] while, forexternal low speed flows, the reference temperature practically coincides with thestream one.

From Eqs. (1)–(3) it is possible to find an explicit expression for h:

h ¼Qj � s�ðT4

w � T4ambÞ

Tw � T r. (4)

Under the assumption that the Biot number Bi ¼ hs=l (where s and l are thethickness and thermal conductivity coefficient of the foil, respectively) is small ascompared to unity, temperature can be considered practically constant across the foilthickness. Therefore, the surface of the foil to be measured can also be chosen as thatopposite to the heat exchange surface.

If this surface is not fully adiabatic (see Fig. 1b), Eq. (1) should be extended toinclude the total heat flux to external ambient Qa. Usually, this heat flux results to bethe sum of the radiative and the natural convection heat fluxes. The naturalconvection heat flux to external ambient can be evaluated by using standardcorrelations tables [15–17] or, better, by making some ad hoc tests.

The hypothesis of zero-dimensional sensor is rigorously satisfied only if theconstant heat generation over the sensor surface leads to a spatially constanttemperature of the sensor itself, i.e. practically when the convective heat transfercoefficient is constant too. However, in many thermo-fluid-dynamic phenomen-ologies the heat transfer coefficient varies and this involves variations of the sensorsurface temperature as well. These variations cause conductive heat fluxes in thetangential (to the sensor surface) direction, which may constitute an important partof the total heat flux (Fig. 1c). By retaining the assumption that the sensor isthermally thin (i.e. with a constant temperature across its thickness) and ideal, it ispossible (for an isotropic slab) to evaluate the tangential conduction heat flux Qk bymeans of Fourier law:

Qk ¼ �lsr2Tw. (5)

Therefore, in order to extend the heated-thin-foil technique to the multi-dimensional case it is necessary to include in the energy balance the conductiveheat flux along the tangential direction as well. So the final form of the convectiveheat transfer coefficient becomes

h ¼Qj � s�ðT4

w � T4ambÞ �Qa þQk

Tw � T r. (6)

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It is important to remark that the use of IRSR (intrinsically two-dimensional)generally enables to evaluate the Laplacian of Eq. (5) by numerical computation. Ofcourse, this can be performed only after an adequate filtering of the cameraexperimental signal, which is typically affected by noise.

In many applications of the heated-thin-foil sensor, a quasi spatially constant Jouleheating can be easily obtained by using a printed circuit board [18,19]. The printedcircuit is generally manufactured by several adjacent thin (down to 5 mm) coppertracks arranged in a greek fret mode (see Fig. 2) and bound to a fibreglass substrate.Due to the high conductivity coefficient of copper, this printed circuit board has ananisotropic thermal conduction behaviour (along or across the tracks) so that it isnot possible to evaluate the conductive heat flux by means of the classical Fourierlaw (5). By still retaining the assumption that Tw is independent of the coordinate z

which is normal to the slab, it is therefore necessary to generalise Eq. (5) to take intoaccount this effect:

Qkðx; yÞ ¼ �rðsðx; yÞLðx; yÞ rTwðx; yÞÞ. (7)

To simplify Eq. (7), it is feasible to roughly separate the effect due to the coppertracks from that of the fibreglass support. In particular, by choosing a Cartesiancoordinate system with its axes directed as the two principal axes of the thermalconductive tensor L (see Fig. 2), it is possible to split the effects in the directionsnormal and parallel to the copper tracks [19,20]. In this case, the total conductiveheat flux may be expressed as the sum of two contributions one along the x directionQkx and the other one Qky along the y-axis:

Qk ¼ Qkx þQky. (8)

Fig. 2. Printed circuit board.

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Bearing in mind the sketch of Fig. 2, it is easy to understand that along the y-axis,the conductive heat flux is the sum of two mechanisms in parallel, one due to thecopper tracks and the other one to the fibreglass support. By considering the meanheat flux, it is obtained:

Qky ¼ �wcsclc þ wfsflf

wf

� �q2Tw

qy2

¼ � ðg�sclc þ sflf Þq2Tw

qy2¼ �ðslÞey

q2Tw

qy2, ð9Þ

where w indicates width; s thickness; the suffixes c and f are relative to copper and tofibreglass, respectively, and it has been introduced the width parameter g� defined as

g� ¼wc

wf. (10)

In Eq. (9), the quantity ðslÞey stands for the equivalent thermal conductance alongthe y-axis while wf represents also the greek pitch.

The phenomenon is slightly more complicated in the direction normal to thecopper tracks. In fact, in the copper gap only fibreglass allows conductive heattransfer while, in the track zone, both materials contribute to it. Therefore, in thiscase, the conductive heat transfer can be estimated as due to both a series and aparallel processes:

Qkx ¼ �1� g�

sflfþ

g�

sclc þ sflf

� ��1 q2Tw

qx2¼ �ðslÞex

q2Tw

qx2, (11)

where (sl)ex represents the equivalent conductance along the x-axis.As expected in the limits g� ! 0, or g� ! 1, both Eqs. (9) and (11) reduce to the

case of an isotropic material. For the typical case of sclc=sflf ¼ 17, Fig. 3 shows theequivalent conductances, referred to that of fibreglass, in both the direction of the

Fig. 3. Equivalent thermal conductance (sclc=sflf ¼ 17).

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copper tracks and that normal to them. For g� less than 0.8, the equivalentconductance in the direction orthogonal to the copper tracks, reduces to less thanone fourth of the value relative to g� ¼ 1. This event may be exploited whenever thepreferred direction of the spatial temperature gradient is a priori known to reducetangential conduction.

4. Application of IRSR to transient techniques

As already pointed out, IRSR can be regarded as a two-dimensional array of thinfilms. In the transient technique, however, the measured temperatures can becorrelated to the heat flux by using either the one-dimensional semi-infinite wallmodel, or the wall calorimeter (Figs. 4 and 5). In the former case, practically theheat-flux sensor will be anyhow constituted by a slab of finite thickness s; hence thethin-film model may be applicable only for relatively small measurement times (i.e.,there is a lower limit to the frequencies the sensor gives trustworthy results). On aquantitative basis, if tM is the measuring time, it has to be verified:

tMos2

2a, (12)

where a is the slab thermal diffusivity coefficient. Therefore, for this sensor theboundary condition on the other surface is irrelevant as long as the assumption ofsemi-infinite wall is valid.

By assuming the thin-film sensor to be isothermal at initial time t ¼ 0, a suitableformula to evaluate the heat flux from the measured surface temperature is [21]

Qc �Qr ¼

ffiffiffiffiffiffiffiffirclp

rfðtÞffiffi

tp þ

1

2

Z t

0

fðtÞ � fðxÞ

ðt� xÞ3=2dx

" #, (13)

where f ¼ TwðtÞ � Twi is the surface temperature difference (Twi being the initialvalue of the wall temperature Twi ¼ Twð0Þ); r, c and l, are the mass density, the

Fig. 4. Thin-film sensor.

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Fig. 5. Thin-skin sensor.

T. Astarita et al. / Optics and Lasers in Engineering 44 (2006) 261–281 269

specific heat and the thermal conductivity coefficient of the sensor material,respectively.

Usually, the integral of Eq. (13) is numerically evaluated by using one of thealgorithms accepted for aerospace applications [22]. However, such algorithms aregenerally sensitive to temperature measurement errors and one should be verycautious when using them with noisy data and/or when the initial time is notprecisely known. Moreover, the approach based on Eq. (13) needs a relatively highdata sampling rate and this requirement is not often fully satisfied by standardIRSRs due to their maximum acquisition frequency typically of the order of 50Hz.

An alternative approach [23], that works much better in these cases, is based onthe assumption that the direct problem yields a certain heat-flux time variation law,where some free parameters are present. Then such parameters are found so that thecomputed temperatures best agree with the experimentally measured temperatures.The best fit may be determined by the ordinary least squares criterion.

In the most common case of a constant heat transfer coefficient h and constantreference temperature T r, the convective heat transfer rate varies linearly with thewall over-temperature. Based on the above boundary condition, the solution of theheat diffusion equation in solids can be obtained by Laplace transforms as

Tw ¼ Twi þ ðT r � TwiÞð1� eb2

erf cbÞ (14)

with b ¼ hffiffitp=ffiffiffiffiffiffiffiffircl

p.

In the presence of a radiative heat flux and under the assumption that theconvective and radiative contributions are uncoupled, Eq. (14) may be modified totake into account the radiative correction:

Tw ¼ Twi þ ðT r � TwiÞð1� eb2

erf cðbÞÞ �Qr

h. (15)

The least-squares method consists of finding h and Twi (which, to a certain extent,may be not correctly determined due to inaccuracy on temperature measurementand/or on starting time) to minimise the function

Xn

j¼1

ðY j � TwjÞ2, (16)

where Y j is the j-term of the n experimentally measured surface temperature valuesand Twj is the temperature predicted by means of Eq. (15). Both of thesetemperatures are evaluated at the same time and at the same location.

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In the case of the wall calorimeter (thin-skin), the sensor, practically a thin plate(Fig. 5), is modelled as an ideal calorimeter (isothermal across its thickness) which isheated on one surface and thermically insulated on the other one. An unsteady one-dimensional energy balance gives

Qc þQr ¼ rcsdTw

dt, (17)

where Tw is the sensor temperature.From Newton law, Eq. (3) and by knowing the temperature evolution (to be

measured with the IRSR), it is possible to evaluate the convective heat transfercoefficient. The use of IRSR in the wall calorimeter technique is quite advantageousbecause the temperature can be measured on either side of the model.

As already mentioned, for both thin-skin and thin-film models, the heat flux withinthe sensor is generally assumed to be one-dimensional. This hypothesis is rigorouslysatisfied only when the temperature over the sensor surface is constant. However, inmany thermo-fluid-dynamic phenomenologies, the involved heat flux (and corre-spondingly the temperature) varies over the surface. Under the assumption that thesensor material is isotropic, or (as already done in the previous paragraph) bychoosing a cartesian coordinate system with its axes directed as the two principalaxes of the thermal conductive tensor, it is possible to split conduction effects in thetwo tangential directions. For the sake of ease, in the following, it is assumed that theconvective heat flux harmonically varies only along one direction parallel to thesensor surface the extension to any arbitrary convective heat flux being straightfor-ward.

A suitable expression for steady convective heat fluxes harmonically varying in thex direction is the following:

QcðxÞ ¼ Qu þQh cosðkxÞ, (18)

where Qu represents the steady part of the heat flux, Qh is the amplitude of itsharmonic part, and k ¼ 2p=L is the wave number (L being the wavelength).

For the two sensors, the response due to a harmonic spatial variation of the heatflux is given by de Felice et al. [24] in terms of difference between the surfacetemperature Twðx; tÞ at time t and the initial uniform temperature Twi:yðx; tÞ ¼ Twðx; tÞ � Twi.

For both sensors it results:

yðx; tÞ ¼ Bf ðFoÞ cosðkxÞ, (19)

where Fo ¼ k2at.If suffix t denotes the thin-skin sensor and suffix m the thin-film one, it is

Bt ¼ Qh=ðlk2sÞ; f t ¼ 1� exp ð�FoÞ, (20)

Bm ¼ Qh=ðlkÞ; f m ¼ erfðffiffiffiffiffiffiFopÞ, (21)

where s is the thickness of the thin-skin sensor.Eq. (19) states that, in both cases, there is no phase difference between the

incident harmonic heat flux and the surface temperature response. The maximum

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amplitudes, obtained for Fo!1, are Bt and Bm, respectively. For finite values ofFo, they are reduced by the attenuation factors f t and f m, respectively.

To correct the measured temperatures so as to take into account the tangentialconduction effects, it is convenient to evaluate the ratio between the temperatureamplitude B f(Fo) (as given by Eqs. (20) and (21)) and that corresponding to thesame value of Qh but in absence of tangential conduction (which is given by theclassical one-dimensional solutions). By defining this ratio as temperature amplitude

transfer function (A), for the two models it results:

At ¼1� exp ð�FoÞ

Fo(22)

and

Am ¼

ffiffiffipp

2

erfffiffiffiffiffiffiFopffiffiffiffiffiffiFop . (23)

The amplitude of each harmonic component of the measured temperature may bethus corrected and the corresponding harmonic component of the heat flux can beevaluated by using the classical one-dimensional formulae. Af and Am are plotted asa function of the Fourier’s number in Fig. 6 which shows that the thin-film sensor hasto be generally preferred to the thin-skin one because of its lower modulation oftemperature amplitude. However, being s=L51, the ratio of the temperaturemaximum amplitude is favourable to the thin-skin sensor.

Fig. 6. Temperature amplitude transfer function.

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5. Applications

In the following, the heat transfer in three different fluid flow configurations areanalysed by means of infrared thermography by using both the steady-state heated-

thin-foil sensor and the unsteady thin-film one. With the former sensor a circularcylinder in a wind tunnel and a 1801 turn channel with and without V rib turbulators(i.e. an external flow and an internal one) are investigated, while the thin-film sensorhas been applied to the study of the shock wave/boundary layer interaction on a flatplate with a ramp in a high enthalpy hypersonic wind tunnel.

5.1. Circular cylinder

Cylindrical bodies with circular cross section placed in a longitudinal flow arefound in many engineering applications; the flow field around them is characterisedby different types and extent of flow separation and reattachment according to thegeometry of the cylinder upstream end and of the angle of attack of their axis relativeto the incoming flow.

The tested longitudinal cylinder has an outer diameter D ¼ 40mm, an overallstreamwise length of 300mm and its lateral surface is made out of a printed circuitboard (bonded to a fibreglass layer) so as to generate a constant Joule heat flux overit. The copper conducting tracks of the printed circuit are 35 mm thick, 3mm wide,placed at 4mm pitch and aligned perpendicularly to the cylinder axis. Two differentconfigurations of the cylinder leading edge (nose) are tested: a sharp edge bluff noseand a hemispherical (round) blunt one.

Tests are performed in an open circuit wind tunnel having a 300� 400mm2

rectangular test section which is 1.1m long. The freestream turbulence intensity ofthe tunnel is quite low and lies in the range 0.08–0.12% depending on the testingconditions. The access window for the infrared camera to the test section of the windtunnel is made of bioriented polyethylene; calibration of the radiometer takes intoaccount its presence.

The convective heat transfer coefficient is calculated by means of Eq. (6), where,because of the stream low Mach number, the adiabatic wall temperature is assumedto coincide with the free stream temperature Taw ¼ T1. Tests are carried out forvarying the Reynolds number Re (based on the diameter of the cylinder D and on thefreestream velocity V1) from 26,000 to 89,000 and the angle of attack of the cylinderaxis with respect to the oncoming flow g from 01 to 101.

In order to measure temperatures in the whole heated zone and to account for thedirectional emissivity coefficient, three thermal images in the azimuthal direction aretaken and patched up. In particular, to reduce the measurement noise, each image isobtained by averaging 32 thermograms in a time sequence. It has to be noted that,due to the end-conduction effects near the forebody, the portion of the cylinder forwhich the infrared camera gives reliable data actually starts at x=D ¼ 0:2 (x beingthe coordinate along the cylinder axis) and data are reported up to x=D ¼ 5. Thiszone is precisely identified by putting markers over the cylinder surface, which areuseful also to patch up the various thermal images.

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The flow field around a cylindrical body is characterised by separation andreattachment of the flow, which can be inferred from the distribution of the heattransfer coefficients. The heat transfer coefficients are computed in non-dimensionalform in terms of the Nusselt number Nu based on cylinder diameter. It must beremembered that, as proved by Sparrow et al. [25], the location of the maximum Nu

does not exactly coincide with that of flow reattachment; however, the position ofthe maximum Nu can be considered to determine the length of the thermalseparation bubble [26].

For g ¼ 01 the maximum Nu value is positioned at x=D ¼ 1:621:7 for the sharpleading edge and does not depend on Re. Instead for the round nose, the maximumNu value position depends strongly on Re since it moves from x=D ¼ 0:3 to 0.7 asRe decreases from 89,000 to 26,000. Results of the present investigation confirmthe assertions of Carlomagno [27,28] about the fundamental role played by thefreestream turbulence level for the formation of the leading edge separationbubble.

As g increases from 01 to 101, for both configurations (sharp and round), themaximum Nu moves upstream on the windward side while it remains quite in thesame position on the leeward one. For Re ¼ 71; 000 and g ¼ 101 some of theobtained data are presented in terms of Nu isocontours in Fig. 7 for the sharp edgeand in Fig. 8 for the round nose. As it can be seen in Fig. 7 (sharp edged cylinder),the separation bubble appears shorter on the windward side, with respect to that onthe leeward one, and at reattachment the Nusselt number assumes also higher values.On the contrary, for the round nosed cylinder (Fig. 8) two thermal reattachmentpoints are present on the leeward side. A likely explanation for this is thatthe separation bubble disappears on the windward side giving rise to theformation of two vortices, which can be assumed to coincide with the saddle pointsobserved by Peake and Tobak [29] on either side of the nodal separation point on theleeward side.

Fig. 7. Nu isocontours for sharp-edged cylinder, Re ¼ 71; 000, g ¼ 101.

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Fig. 8. Nu isocontours for round-nosed cylinder, Re ¼ 71; 000, g ¼ 101.

T. Astarita et al. / Optics and Lasers in Engineering 44 (2006) 261–281274

Another feature, more evident for the round nose cylinder, is the appearanceof a low heat transfer region on the cylinder sides. The latter, by increasing theangle of attack, moves upstream simultaneously becoming sharper andenhancing the three-dimensionality of the flow. This region is presumably connectedwith the fact that the increasingly intense cross-flow leads first to instabilities of theboundary layer, and eventually to the separation from the sides of the cylinder ofdominating longitudinal vertical structures, similar to those described by Peake andTobak [29].

5.2. The 1801 turn channel with and without V ribs

This flow configuration is often encountered inside turbine blades for coolingpurposes. Really rib turbulators are often also used in the design of heat exchangerchannels in order to enhance the convective heat transfer rate and thus allowing toboth reducing the overall exchanger dimensions and to increase efficiency. In 1801turn channels, the flow is quite complex due to the various separations andreattachments of the flow and this behaviour it is further enhanced in the presence ofrib turbulators.

A two-pass channel of square cross-section 80� 80mm2 and 2000mm long beforethe turn is tested; these dimensions guarantee a hydro-dynamically fully developedflow ahead of the 1801 turn. The central partition wall between the two adjacentducts is 16mm thick and ends with a square tip 80mm distant from the short endwall of the channel. The two side walls of the channels are heated by means of threeprinted circuit boards and square rib turbulators (8mm in side), made of aluminium(to have a high thermal conductance), are glued to them. Ribs have a V shape, withan angle of 451 with respect to the duct axis, have their apex pointing downstreamand are placed at a rib-pitch to rib-side ratio P/e of either 10 or 20. Further detailsabout the experimental apparatus can be found in [30].

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The heat transfer coefficient is calculated by means of Eq. (6) where Tr coincideswith the local bulk temperature Tb which is evaluated by measuring the stagnationtemperature at the channel entrance and by making a one-dimensional energybalance along the channel. Data are reduced in non-dimensional form in terms of theNusselt number normalised by its fully developed counterpart Nu� (Dittus–Boeltercorrelation [31]). Both the Nusselt number Nu and the Reynolds number Re arebased on the channel hydraulic diameter.

For the smooth channel and Re ¼ 30; 000, the distribution of the local Nu=Nu� inthe vicinity of the turn, is reported in Fig. 9a. Air enters the channel from the lowerduct and leaves from the upper one. By moving streamwise along the channel, theratio Nu=Nu� increases around the turn and downstream of it because of thepresence of secondary flows. Three relatively high heat transfer regions may berecognised: the first one is located by the end wall (in front of the partition walltowards the first outer corner) and is caused by the jet coming from the first ductwhich impinges on this wall; the second one is located at the outer wall downstreamof the second corner and is due to the jet effect of the flow through the bend; thethird one is located at about the half part of the partition wall, downstream of thesecond inner corner, where the flow rebounding from the outer wall, impinges beforeexhausting. The second zone attains Nu=Nu� values much greater than the other twodue to the presence of strong secondary flows already found by Arts et al. [32]. Tworelatively low heat transfer zones are also observed, one just before the first corner ofthe outer wall and the other one in the neighbourhood of the tip of the partition wall;these zones constitute evidence for the existence of recirculation patterns.

The overall increase of the convective heat transfer coefficient due to the presenceof ribs is clearly evident from Fig. 9b and c, where Nu=Nu� are shown for the twotested rib-pitch to rib-side ratios P/e. In all the Nu maps ribs are clearly visible due tothe higher heat transfer rate that occurs on them. For both dimensionless pitches, thethermal pattern before the turn appears to be repetitive (in a sense, the flow could beconsidered as thermally fully developed). For example, in Fig. 9c, the shape and levelsof the contour lines after the first three ribs of the inlet duct are practically identical.Instead, some differences due to some measurements edge effects are found at theduct entrance.

The secondary flows induced by the V-shaped ribs have the form of two pair ofcounter rotating cells and produce variations in the spanwise Nusselt numberdistribution both in the inlet and in the outlet channel by decreasing the convectiveheat transfer coefficient towards the channel axis with respect to that nearby thewalls.

Especially in the inlet duct, the reattachment line downstream of the ribs can beidentified by the locus of the normalised Nusselt number local maxima when movingin the mean streamwise direction. The reattachment distance, which increases forthe higher rib pitch, appears also to increase going from the walls towards thechannel axis and this is most likely due to the interaction of the main flow with thesecondary one.

In the proximity of the first external corner, it is possible to see a low heat transferzone, due to a recirculation bubble as already observed for the smooth channel. Just

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Fig. 9. Normalised Nusselt number isocontours for Re ¼ 30; 000. (a) Smooth, (b) P=e ¼ 20, (c) P=e ¼ 10.

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after the last rib and in proximity of the partition wall, the interaction between thesecondary flow and the sharp turn produces a high heat transfer zone that tends toshift downstream for increasing pitch.

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For both pitches, the overall increase of the turbulence due to the bend induceshigher values of the normalised Nusselt number in the outlet duct but the percentageincrease is quite lower than what occurring in the smooth channel.

5.3. Shock-wave/boundary-layer interaction

The development of hypersonic vehicles has renewed the attention on the problemof viscous inviscid flow interactions and, in particular, on shock-wave/boundary-layer interaction phenomena that are of great practical importance for air-breathingengine inlets, wing/body junctures and deflected control surfaces. Prediction ofthermal and dynamic loads on surfaces exposed to hypersonic flows is an essentialprerequisite for the effective design of aerodynamic control surfaces and of thermalprotection system of modern space vehicles in their trans-atmospheric flight portion.

Measurements presented in this section refer to shock wave-boundary layerinteraction in a two-dimensional hypersonic flow over a model consisting of a flatplate followed by a compression ramp (wedge) with its hinge line parallel to themodel’s leading edge.

The model surface is realised with two separate MACORTM plates screwed ontoaluminium supports. The model spanwise dimension is 100mm. The hinge line ispositioned at 50mm from the leading edge and the ramp angle is 151. MACORTM

was chosen as the model surface material for its low thermal conductivity, asrequired in connection with the use of thin film model.

Experimental tests have been carried out in Centrospazio high-enthalpy arc-heated tunnel (HEAT) [33,34] that is capable of producing Mach 6 flows with aspecific total enthalpy up to 2.5MJ/kg on an effective test section 60mm in diameter,in the low to medium Reynolds number range (104–106). The tunnel operates in apulsed, quasi-steady mode, with running time ranging from 50 to 200ms. HEATfacility mainly consists of an arc gas heater and a contoured expansion nozzle,installed in a vacuum chamber volume of 4.1m3; auxiliary systems are fitted to thearc heater to provide it with working fluid and energy. Four rotary pumps evacuatethe chamber until an ultimate pressure of 10 Pa is reached before each run. Thisvacuum level allows an under-expanded hypersonic flow-field to be maintained at thenozzle exit for a running time longer than 200ms. IR camera used during the test wasFLIR SC 3000 and acquisition frame frequencies was 60Hz for flow visualisationand 300Hz for heat-flux measurements.

A thermal map recorded about 80ms after the starting of wind tunnel is reportedin Fig. 10. The temperature distribution is almost bidimensional only near the modelleading edge (the flow comes from left to right). Moving downwind, the continuousdecrease of wall temperature, indicates the development of the boundary layer. Nearthe hinge line is clearly visible a region where the temperature attains a minimumthat is due to the presence of a separation region in the flow. Moving along thesymmetry axis after the hinge line the temperature reaches a maximum that is to becorrelated to the flow reattachment on the ramp.

If one considers that, in a first approximation, the potential core may beassimilated to a cone emerging from the nozzle exit (the cone height being

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Fig. 10. Temperature map (in 1C) recorded on model surface after 80ms from tunnel starting. Total

enthalpy: 2.3MJ/kg; stagnation pressure: 4.6 bar.

Fig. 11. Stanton number profile on symmetry axis. Total enthalpy: 1.8MJ/kg; stagnation pressure: 6 bar.

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determined by the expansion fan angle at the nozzle exit, ideally starting at arcsin (1/M)), the intersection of this cone with model surface is clearly visible on the thermalmap. The measured temperature time histories are used to compute heat flux with thin-film model described in Section 4. In this case, it was not possible to use the alternativeapproach proposed in [23] because during the first 30ms of test run the total enthalpy(and therefore the reference temperature) is not constant. For two typical runs, in Fig.11 the heat flux along the symmetry axis is presented in non-dimensional form bymeans of the Stanton number based on the adiabatic wall temperature computed bymeans of the recovery factor for laminar boundary layer flow [23].

Experimental data are also compared with the classical flat plate boundary layeranalytical solution [35]. The results show a good agreement with theoretical solutionon the flat plate. Near the hinge line (X ¼ 50mm) the presence of a separation regionis clearly identified from the minimum of the Stanton number distribution. The

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entity of the heat flux at reattachment is in good agreement with data present inliterature.

6. Conclusions

The application of infrared thermography as an optical method in heat transferand fluid flow visualisation is analysed. The heat-flux sensors, which are normallyused for the measurements of convective heat transfer coefficients, and theapplication of the infrared scanning radiometer as a temperature measuring deviceare critically reviewed. In particular, the corrections of the errors associated withtangential conduction along the sensor are investigated for the heated-thin-foil, thethin-film and the wall calorimeter sensors.

The heated-thin-foil heat flux sensor coupled with measurement of surfacetemperature by IR thermography is used to measure the convective heat transfercoefficient on two flowfields: a circular cylinder at an angle of attack and a 1801 turnchannel with and without rib turbulators. Furthermore the thin-film sensor has beenapplied to the study of the shock wave/boundary layer interaction in a flat plate witha ramp in a high enthalpy hypersonic wind tunnel.

For the circular cylinder, in order to measure temperatures, in the whole heatedzone and to account for the directional emissivity coefficient, three thermal images inthe azimuthal direction are taken and patched up. For the sharp edge cylinder andan angle of attack of 101, the separation bubble appears shorter on the windwardside, with respect to that on the leeward one, and at reattachment the Nusseltnumber assumes also higher values. On the contrary, for the round nosed cylindertwo thermal reattachment points are present on the leeward side, while noreattachment is evident at the windward one.

In the inlet zone, ribbed channels show spanwise variations of the heat transfermaps because of the presence of secondary flows. For both tested rib pitches, theoverall increase of turbulence due to the bend induces higher values of thenormalised Nusselt number, but, in the outlet duct, the percentage increase is lowerthan that relative to a smooth channel because of the rib already induced turbulence.This should decrease the thermal stresses in the turbine blade.

Shock-wave/boundary-layer interaction phenomena in high enthalpy hypersonicflows has been studied by means of heat-flux measurement performed by IRthermography coupled with thin-film sensor. The use of IR thermographydemonstrate that the flow condition are two dimensional only geometrically.However, on the symmetry axis the IR quantitative measurements are in goodagreement with literature data.

References

[1] Gardon R. A transducer for the measurement of heat flow rate. Trans J Heat Transfer 1960;82:396–8.

[2] Vidal RJ. Transient surface temperature measurements. CAL Rep 1962;114:1–55.

ARTICLE IN PRESS

T. Astarita et al. / Optics and Lasers in Engineering 44 (2006) 261–281280

[3] Scott CJ. Transient experimental techniques for surface heat flux rates. Measurements Techniques in

Heat Transfer, AGARDograph, Vol. 130, 1970. p. 309–28.

[4] Willeke K, Bershader D. An improved thin-film gauge for shock-tube thermal studies. Rev Sci

Instrum 1973;44:22–5.

[5] Baines DJ. Selecting unsteady heat flux sensors. Instrum Control Syst 1972:80–3.

[6] Thompson WP. Heat transfer gages. In: Marton L, Marton C, editors. Methods of experimental

physics, vol. 18B. New York: Academic; 1981. p. 663–85.

[7] Balageas DL, Boscher DM, Deom AA, Fournier J, Gardette G. Measurement of convective heat-

transfer coefficients in wind tunnels using passive and stimulated infrared thermography. Rech

Aerosp 1991;4:51–72.

[8] Carlomagno GM. Thermo-fluid-dynamics applications of quantitative infrared thermography.

J Flow Visualization Image Process 1997;4:261–80.

[9] De Luca L, Cardone G, Carlomagno GM, Aymer D, Alziary T. Flow visualization and heat transfer

measurements in hypersonic wind tunnel. Exp Heat Transfer 1992;5:65–79.

[10] Nortershauser D, Millan P. Resolution of a three-dimensional unsteady inverse problem by

sequential method using parameter reduction and infrared thermography measurements. Numer

Heat Transfer Part A 2000;37:587–611.

[11] Shapiro AH. The dynamics and thermodynamics of compressible fluid flow, Vols. I and II. New

York: Ronald Press; 1954.

[12] Zucrow MJ, Hoffman JD. Gas dynamics, vols. I and II. New York: Wiley; 1976.

[13] Owczareck JA. Gas dynamics. International Textbook Company; 1964.

[14] Meola C, de Luca L, Carlomagno GM. Influence of shear layer dynamics on impingement heat

transfer. Exp Thermal Fluid Sci 1996;13:29–37.

[15] Kays WM, Crawford ME. Convective heat and mass transfer. New York: McGraw-Hill; 1993.

[16] Perry JH. Chemical engineers’ handbook. New York: McGraw-Hill; 1963.

[17] Kakac S, Shah RK, Aung W. Handbook of single phase flow convective heat transfer. New York:

Wiley; 1987.

[18] Cardone G, Astarita T, Carlomagno GM. Heat transfer measurements on a rotating disk in still air.

Proceedings of Flucome’94, Toulouse. Vol. 2, 1994. p. 775–80.

[19] Astarita T. Alcuni aspetti di scambio termico nelle turbine a gas, PhD thesis. University of Naples,

1996.

[20] Astarita T, Cardone G. Thermofluidynamic analysis of the flow near a sharp 1801 turn channel. Exp

Thermal Fluid Sci 2000;20:188–200.

[21] Baines DJ. Selecting unsteady heat flux sensors. Instr Control Syst 1972:80–3.

[22] Cook WJ, Felderman EJ. Reduction of data from thin-film heat-transfer gages: a concise numerical

technique. AIAA J 1966;4:561–2.

[23] de Luca L, Cardone G, Aymer de la Chevalerie D, Fonteneau A. Experimental analysis of viscous

interaction in hypersonic wedge flow. AIAA J 1995;33(12):2293–8.

[24] de Felice G, de Luca L, Carlomagno GM. La misura dei flussi termici convettivi nel caso di

distribuzioni non uniformi. Proceedings of the VII congr Naz UIT, Firenze. 1989. p. 591–9.

[25] Sparrow EM, Kang SS, Chuck W. Relation between the points of flow reattachment and maximum

heat transfer for regions of flow separation. Int J Heat Mass Transfer 1987;30:1237–46.

[26] Cardone G, Buresti G, Carlomagno GM. Heat transfer to air from a yawed circular cylinder. In:

Nakayama Y, Tanida Y, editors. Atlas of visualization III. Boca Raton, FL: CRC Press; 1997.

p. 153–68 [Chapter 10].

[27] Carlomagno GM. Heat transfer measurements and flow visualization performed by means of infrared

thermography. In: Di Marco P, editor. Proceedings of eurotherm seminar 46. Heat transfer in single

phase flows, Vol. 4. Pisa; 1995. p. 45–52.

[28] Carlomagno GM. Quantitative infrared thermography in heat and fluid flow. Optical methods and

data processing in heat and fluid flow. IMechE Conference Transactions. Vol. 3. London; 1996.

p. 279–90.

[29] Peake DJ, Tobak M. Three-dimensional flows about simple components at angle of attack.

AGARD-LS-121, Paper 2, 1982.

ARTICLE IN PRESS

T. Astarita et al. / Optics and Lasers in Engineering 44 (2006) 261–281 281

[30] Astarita T, Cardone G, Carlomagno GM. Convective heat transfer in ribbed channels with a 1801

turn. Exp Fluids 2002;33:90–100.

[31] Dittus PW, Boelter LMK. Heat transfer in automobile radiators of the tubular type. Univ Calif Pub

Eng 1930;2(13):443–61 (reprinted in Int J Comm Heat Mass Transfer 1985;12:3–22).

[32] Arts T, Lambert de Rouvroit M, Rau G, Acton P. Aero-thermal investigation of the flow developing

in a 180 degree turn channel. Proceedings of the International Symposium on Heat Transfer in

Turbomachinery. Athens, 1992.

[33] Scortecci F, Paganucci F, d’Agostino L, Andrenucci M. A new hypersonic high enthalpy wind tunnel.

The 33rd joint propulsion conference, AIAA 97-3017. Seattle, 1997.

[34] Scortecci F, Paganucci F, Biagioni L. Development of a pulsed arc-heater for a hypersonic high

enthalpy wind tunnel. The 33rd joint propulsion conference. AIAA 97-3016, Seattle, 1997.

[35] Simeonides G. Hypersonic shock wave boundary layer interactions over compression corners. PhD

thesis, Dip of Aerospace Engineering, Faculty of Engineering, University of Bristol, 1992.