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Transcript of Basic Thermography
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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. AstaritaÃ, G. Cardone, G.M. Carlomagno
Universita degli studi di Napoli ‘‘Federico II’’, Dipartimento di Energetica Termofluidodinamica Applicatae 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
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0143-8166/$ - see front matterr 2005 Elsevier Ltd. All rights reserved.
doi:10.1016/j.optlaseng.2005.04.006
ÃCorresponding author. Tel.: +39 081 768 3389; fax: +39 081 239 0364.
E-mail address: [email protected] (T. Astarita).
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temperature detectors or pyrometers, each transducer yields either the heat flux at a
single point, or the space-averaged one; hence, in terms of spatial resolution, the
sensor itself can be considered as zero-dimensional. This constraint makes
experimental 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 infrared
scanning radiometer (IRSR) constitutes a true two-dimensional temperature
transducer since it allows the performance of accurate measurement of surface
temperature maps even in the presence of relatively high spatial temperature
variations. Correspondingly, the heat-flux sensor may become two-dimensional as
well. In particular, infrared thermography can be fruitfully employed to measure
convective heat fluxes, in both steady and transient techniques [7–9]. Within this
context, IRSR can be intrinsically considered as a thin-film sensor [5] because it
generally measures skin temperatures. The thermal map obtained by means of
currently available computerised thermographic systems is formed through a large
amount of pixels (20–300 K and more) so that IRSR can be practically regarded as a
two-dimensional array of thin films. However, unlike standard thin films, which have
a response time of the order of microseconds, the typical response time of IRSR is of
the order of 10À1 –10À3 s.
The use of IRSR as a temperature transducer in convective heat transfer
measurement appears, from several points of view, advantageous if compared to
standard 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 the
conduction errors through the thermocouple or resistance temperature detector
wires.
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 heat
conduction 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 element
surface, i.e. the temperature gradient components that are parallel to the slab plane
are neglected. In practice, the slab surfaces can also be curved, but their curvature
can be ignored if the layer affected by the input heat flux is relatively small as
compared 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-dimensional
case. The term ideal means that thermophysical properties of the sensor material areassumed to be independent of temperature and that the influence of the temperature
sensing 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 measures
the 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 surface
film must be very thin so as to have negligible heat capacity and thermal
resistance as compared to the slab ones. To use this sensor with infrared
thermography, 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 the
time rate of change of the mean slab temperature. This temperature is usually
measured by using the slab as a resistance thermometer.
(3) Wall calorimeter or thin-skin sensor: The slab is made thermally thin (so that its
temperature can be assumed to be constant across its thickness) and, as in the
case of the thick-film sensor, is used as a calorimeter. Heat flux is typically
inferred from the time rate of change of the slab temperature which is usually
measured 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 slab
thickness is measured. By considering a steady-state heat transfer process, heat
flux is computed by means of the temperature gradient across the slab. The
temperature 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 heat
transfer coefficient from an overall energy balance. Also, in this case, due to the
thinness 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 circular
Gardon gauge, in which the heat flux normal to the sensor surface is related to a
radial temperature difference, in the direction parallel to the gauge plane [1]. This
sensor is practically of no use in infrared thermography.
Recently, another type of heat-flux sensor based on a three-dimensional unsteady
inverse 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 very
practical, so these sensors will not be herein described. The heated-thin-foil sensor
represents a quasi-steady technique that will be discussed in the next paragraph; the
thin-film and the wall calorimeter sensors constitute transient techniques that will be
treated in the following one.
3. The heated-thin-foil steady-state technique
Within the class of steady-state techniques to measure convective heat fluxes
between a fluid stream and a surface, a method, where the application of IRSR seems
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to be very effective, is the heated-thin-foil technique. The sensor is made of a thin
metallic foil which is heated by Joule effect (see the sketch of Fig. 1a). The main
limitation of this technique is that, for practical reasons, the exchanging surface
should have a cylindrical, or conical, geometry.
In the following, it is initially supposed that the sensor is one-dimensional and that
the surface not exposed to flow is adiabatic. By making a very simple (one-dimensional) steady-state energy balance, it is found
Q j ¼ Qr þ Qc, (1)
where Q j is the imposed constant Joule heating per unit area, Qr is the radiative heat
flux to ambient, and Qc is the convective heat flux to fluid.
The radiative heat flux can be evaluated by
Qr ¼ sðT 4w À T 4ambÞ, (2)
where s is the Stefan–Boltzmann constant, is the total emissivity coefficient, and T w
and T amb are the temperature of the wall and of the experimental ambient,respectively. When standard techniques are used to measure the wall temperature, it
is possible to have a very low wall emissivity coefficient so as to ignore the radiative
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Fig. 1. Heated-thin-foil sensor.
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heat flux to ambient. Obviously, this is not the case when measuring temperatures by
means of IRSR.
The convective heat flux can be expressed according to Newton law:
Qc ¼ hðT w À T rÞ, (3)
where h is the convective heat transfer coefficient and T r is a reference temperature.
The reference temperature depends on the stream experimental conditions. For
example, for high Mach number flows (or for the mixing of two streams at different
temperatures), the correct choice is the adiabatic wall temperature [11–14] while, for
external low speed flows, the reference temperature practically coincides with the
stream one.
From Eqs. (1)–(3) it is possible to find an explicit expression for h:
h ¼ Q j À sðT 4w À T
4ambÞ
T w À T r. (4)
Under the assumption that the Biot number Bi ¼ hs=l (where s and l are the
thickness and thermal conductivity coefficient of the foil, respectively) is small as
compared to unity, temperature can be considered practically constant across the foil
thickness. Therefore, the surface of the foil to be measured can also be chosen as that
opposite to the heat exchange surface.
If this surface is not fully adiabatic (see Fig. 1b), Eq. (1) should be extended to
include the total heat flux to external ambient Qa. Usually, this heat flux results to be
the sum of the radiative and the natural convection heat fluxes. The naturalconvection heat flux to external ambient can be evaluated by using standard
correlations tables [15–17] or, better, by making some ad hoc tests.
The hypothesis of zero-dimensional sensor is rigorously satisfied only if the
constant heat generation over the sensor surface leads to a spatially constant
temperature of the sensor itself, i.e. practically when the convective heat transfer
coefficient is constant too. However, in many thermo-fluid-dynamic phenomen-
ologies the heat transfer coefficient varies and this involves variations of the sensor
surface temperature as well. These variations cause conductive heat fluxes in the
tangential (to the sensor surface) direction, which may constitute an important part
of 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 is
possible (for an isotropic slab) to evaluate the tangential conduction heat flux Qk by
means of Fourier law:
Qk ¼ Àlsr 2T w. (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 conductive
heat flux along the tangential direction as well. So the final form of the convective
heat transfer coefficient becomes
h ¼ Q j À sðT 4w À T 4ambÞ À Qa þ Qk
T w À 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. Of
course, this can be performed only after an adequate filtering of the camera
experimental signal, which is typically affected by noise.In many applications of the heated-thin-foil sensor, a quasi spatially constant Joule
heating can be easily obtained by using a printed circuit board [18,19]. The printed
circuit is generally manufactured by several adjacent thin (down to 5 mm) copper
tracks 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 an
anisotropic thermal conduction behaviour (along or across the tracks) so that it is
not possible to evaluate the conductive heat flux by means of the classical Fourier
law (5). By still retaining the assumption that T w is independent of the coordinate z
which is normal to the slab, it is therefore necessary to generalise Eq. (5) to take into
account this effect:
Qk ðx; yÞ ¼ À r ðsðx; yÞLðx; yÞ r T wðx; yÞÞ. (7)
To simplify Eq. (7), it is feasible to roughly separate the effect due to the copper
tracks from that of the fibreglass support. In particular, by choosing a Cartesian
coordinate system with its axes directed as the two principal axes of the thermal
conductive tensor L (see Fig. 2), it is possible to split the effects in the directions
normal and parallel to the copper tracks [19,20]. In this case, the total conductive
heat flux may be expressed as the sum of two contributions one along the x direction
Qkx and the other one Qky along the y-axis:
Qk ¼ Qkx þ Qky. (8)
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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 the
copper tracks and the other one to the fibreglass support. By considering the mean
heat flux, it is obtained:
Qky ¼ À wcsclc þ wf sf lf
wf
q2T w
q y2
¼ À ðgÃsclc þ sf lf Þ q2T w
q y2¼ ÀðslÞe y
q2T w
q y2, ð9Þ
where w indicates width; s thickness; the suffixes c and f are relative to copper and to
fibreglass, respectively, and it has been introduced the width parameter gà defined as
gà ¼wc
wf . (10)
In Eq. (9), the quantity ðslÞe y stands for the equivalent thermal conductance along
the y-axis while wf represents also the greek pitch.
The phenomenon is slightly more complicated in the direction normal to the
copper tracks. In fact, in the copper gap only fibreglass allows conductive heat
transfer while, in the track zone, both materials contribute to it. Therefore, in this
case, the conductive heat transfer can be estimated as due to both a series and a
parallel processes:
Qkx ¼ À 1 À gÃsf lf
þ gÃsclc þ sf lf
À1
q
2
T wqx2 ¼ ÀðslÞex q
2
T wqx2 , (11)
where (sl)ex represents the equivalent conductance along the x-axis.
As expected in the limits gà ! 0, or gà ! 1, both Eqs. (9) and (11) reduce to the
case of an isotropic material. For the typical case of sclc=sf lf ¼ 17, Fig. 3 shows the
equivalent conductances, referred to that of fibreglass, in both the direction of the
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Fig. 3. Equivalent thermal conductance (sclc=sf lf ¼ 17).
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copper tracks and that normal to them. For gà less than 0.8, the equivalent
conductance in the direction orthogonal to the copper tracks, reduces to less than
one fourth of the value relative to gÃ
¼1. This event may be exploited whenever the
preferred 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 thin
films. In the transient technique, however, the measured temperatures can be
correlated to the heat flux by using either the one-dimensional semi-infinite wall
model, 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 the
thin-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 a
quantitative 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 the
boundary condition on the other surface is irrelevant as long as the assumption of
semi-infinite wall is valid.By assuming the thin-film sensor to be isothermal at initial time t ¼ 0, a suitable
formula to evaluate the heat flux from the measured surface temperature is [21]
Qc À Qr ¼ ffiffiffiffiffiffiffiffircl
p
r fðtÞ ffiffi
tp þ 1
2
Z t
0
fðtÞ À fðxÞðt À xÞ3=2
dx
" #, (13)
where f ¼ T wðtÞ À T wi is the surface temperature difference (T wi being the initial
value of the wall temperature T wi ¼ T wð0Þ); r, c and l, are the mass density, the
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Fig. 4. Thin-film sensor.
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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 the
algorithms accepted for aerospace applications [22]. However, such algorithms are
generally sensitive to temperature measurement errors and one should be very
cautious when using them with noisy data and/or when the initial time is not
precisely known. Moreover, the approach based on Eq. (13) needs a relatively high
data sampling rate and this requirement is not often fully satisfied by standard
IRSRs due to their maximum acquisition frequency typically of the order of 50 Hz.
An alternative approach [23], that works much better in these cases, is based on
the 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 the
computed 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 constant
reference temperature T r, the convective heat transfer rate varies linearly with the
wall over-temperature. Based on the above boundary condition, the solution of the
heat diffusion equation in solids can be obtained by Laplace transforms as
T w ¼ T wi þ ðT r À T wiÞð1 À eb2
erf cbÞ (14)
with b ¼ h ffiffi
tp
= ffiffiffiffiffiffiffiffircl
p .
In the presence of a radiative heat flux and under the assumption that the
convective and radiative contributions are uncoupled, Eq. (14) may be modified to
take into account the radiative correction:
T w ¼ T wi þ ðT r À T wiÞð1 À eb2
erf cðbÞÞ À Qr
h. (15)
The least-squares method consists of finding h and T wi (which, to a certain extent,
may be not correctly determined due to inaccuracy on temperature measurement
and/or on starting time) to minimise the function
Xn
j ¼1
ðY j À T w j Þ2, (16)
where Y j is the j -term of the n experimentally measured surface temperature valuesand T w j is the temperature predicted by means of Eq. (15). Both of these
temperatures are evaluated at the same time and at the same location.
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Fig. 5. Thin-skin sensor.
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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 is
heated on one surface and thermically insulated on the other one. An unsteady one-
dimensional energy balance gives
Qc þ Qr ¼ rcsdT w
dt, (17)
where T w 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 transfer
coefficient. The use of IRSR in the wall calorimeter technique is quite advantageous
because 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 within
the sensor is generally assumed to be one-dimensional. This hypothesis is rigorouslysatisfied only when the temperature over the sensor surface is constant. However, in
many thermo-fluid-dynamic phenomenologies, the involved heat flux (and corre-
spondingly the temperature) varies over the surface. Under the assumption that the
sensor material is isotropic, or (as already done in the previous paragraph) by
choosing a cartesian coordinate system with its axes directed as the two principal
axes of the thermal conductive tensor, it is possible to split conduction effects in the
two tangential directions. For the sake of ease, in the following, it is assumed that the
convective heat flux harmonically varies only along one direction parallel to the
sensor surface the extension to any arbitrary convective heat flux being straightfor-
ward.
A suitable expression for steady convective heat fluxes harmonically varying in the
x direction is the following:
QcðxÞ ¼ Qu þ Qh cosðkxÞ, (18)
where Qu represents the steady part of the heat flux, Qh is the amplitude of its
harmonic 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 heat
flux is given by de Felice et al. [24] in terms of difference between the surface
temperature T wðx; t
Þat time t and the initial uniform temperature T wi:
yðx; tÞ ¼ T wðx; tÞ À T wi.
For both sensors it results:
yðx; tÞ ¼ Bf ðFoÞ cosðkxÞ, (19)
where Fo ¼ k 2at.
If suffix t denotes the thin-skin sensor and suffix m the thin-film one, it is
B t ¼ Qh=ðlk 2sÞ; f t ¼ 1 À exp ðÀFoÞ, (20)
B m ¼ Qh=ðlk Þ; f m ¼ erf ð ffiffiffiffiffiffiFop
Þ, (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 B t and B m, respectively. For finite values of
Fo, 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 tangential
conduction 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 the
same value of Qh but in absence of tangential conduction (which is given by the
classical one-dimensional solutions). By defining this ratio as temperature amplitude
transfer function (A), for the two models it results:
At ¼ 1 À exp ðÀFoÞFo
(22)
and
Am ¼ ffiffiffipp
2
erf ffiffiffiffiffiffi
Fop ffiffiffiffiffiffiFo
p . (23)
The amplitude of each harmonic component of the measured temperature may be
thus corrected and the corresponding harmonic component of the heat flux can be
evaluated by using the classical one-dimensional formulae. Af and Am are plotted as
a function of the Fourier’s number in Fig. 6 which shows that the thin-film sensor has
to be generally preferred to the thin-skin one because of its lower modulation of
temperature amplitude. However, being s=L51, the ratio of the temperature
maximum amplitude is favourable to the thin-skin sensor.
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Fig. 6. Temperature amplitude transfer function.
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5. Applications
In the following, the heat transfer in three different fluid flow configurations are
analysed 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 circular
cylinder 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 sensor
has been applied to the study of the shock wave/boundary layer interaction on a flat
plate 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 are
found in many engineering applications; the flow field around them is characterised
by different types and extent of flow separation and reattachment according to the
geometry of the cylinder upstream end and of the angle of attack of their axis relative
to the incoming flow.
The tested longitudinal cylinder has an outer diameter D ¼ 40 mm, an overall
streamwise length of 300 mm and its lateral surface is made out of a printed circuit
board (bonded to a fibreglass layer) so as to generate a constant Joule heat flux over
it. The copper conducting tracks of the printed circuit are 35 mm thick, 3 mm wide,
placed at 4 mm pitch and aligned perpendicularly to the cylinder axis. Two different
configurations 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 Â 400 mm2
rectangular test section which is 1.1 m long. The freestream turbulence intensity of
the tunnel is quite low and lies in the range 0.08–0.12% depending on the testing
conditions. The access window for the infrared camera to the test section of the wind
tunnel is made of bioriented polyethylene; calibration of the radiometer takes into
account 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 assumed
to coincide with the free stream temperature T aw ¼ T 1. Tests are carried out forvarying the Reynolds number Re (based on the diameter of the cylinder D and on the
freestream velocity V 1) from 26,000 to 89,000 and the angle of attack of the cylinder
axis 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 the
directional emissivity coefficient, three thermal images in the azimuthal direction are
taken and patched up. In particular, to reduce the measurement noise, each image is
obtained 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 for
which the infrared camera gives reliable data actually starts at x=D
¼0:2 (x being
the 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 are
useful also to patch up the various thermal images.
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The flow field around a cylindrical body is characterised by separation and
reattachment of the flow, which can be inferred from the distribution of the heat
transfer coefficients. The heat transfer coefficients are computed in non-dimensional
form 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 of
the maximum Nu can be considered to determine the length of the thermal
separation bubble [26].
For g ¼ 01 the maximum Nu value is positioned at x=D ¼ 1:621:7 for the sharp
leading edge and does not depend on Re. Instead for the round nose, the maximum
Nu value position depends strongly on Re since it moves from x=D ¼ 0:3 to 0.7 as
Re decreases from 89,000 to 26,000. Results of the present investigation confirm
the assertions of Carlomagno [27,28] about the fundamental role played by the
freestream turbulence level for the formation of the leading edge separation
bubble.
As g increases from 01 to 101, for both configurations (sharp and round), the
maximum Nu moves upstream on the windward side while it remains quite in the
same position on the leeward one. For Re ¼ 71; 000 and g ¼ 101 some of the
obtained data are presented in terms of Nu isocontours in Fig. 7 for the sharp edge
and 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 on
the 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 that
the separation bubble disappears on the windward side giving rise to the
formation of two vortices, which can be assumed to coincide with the saddle points
observed by Peake and Tobak [29] on either side of the nodal separation point on the
leeward side.
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Fig. 7. Nu isocontours for sharp-edged cylinder, Re ¼ 71; 000, g ¼ 101.
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Another feature, more evident for the round nose cylinder, is the appearance
of a low heat transfer region on the cylinder sides. The latter, by increasing the
angle of attack, moves upstream simultaneously becoming sharper and
enhancing the three-dimensionality of the flow. This region is presumably connected
with the fact that the increasingly intense cross-flow leads first to instabilities of the
boundary layer, and eventually to the separation from the sides of the cylinder of dominating longitudinal vertical structures, similar to those described by Peake and
Tobak [29].
5.2. The 1801 turn channel with and without V ribs
This flow configuration is often encountered inside turbine blades for cooling
purposes. Really rib turbulators are often also used in the design of heat exchanger
channels in order to enhance the convective heat transfer rate and thus allowing to
both reducing the overall exchanger dimensions and to increase efficiency. In 1801
turn 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 of
rib turbulators.
A two-pass channel of square cross-section 80 Â 80mm2 and 2000 mm long before
the turn is tested; these dimensions guarantee a hydro-dynamically fully developed
flow ahead of the 1801 turn. The central partition wall between the two adjacent
ducts is 16 mm thick and ends with a square tip 80 mm distant from the short end
wall of the channel. The two side walls of the channels are heated by means of three
printed circuit boards and square rib turbulators (8 mm in side), made of aluminium
(to have a high thermal conductance), are glued to them. Ribs have a V shape, with
an 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 details
about the experimental apparatus can be found in [30].
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Fig. 8. Nu isocontours for round-nosed cylinder, Re ¼ 71; 000, g ¼ 101.
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The heat transfer coefficient is calculated by means of Eq. (6) where T r coincides
with the local bulk temperature T b which is evaluated by measuring the stagnation
temperature at the channel entrance and by making a one-dimensional energy
balance along the channel. Data are reduced in non-dimensional form in terms of theNusselt number normalised by its fully developed counterpart Nuà (Dittus–Boelter
correlation [31]). Both the Nusselt number Nu and the Reynolds number Re are
based on the channel hydraulic diameter.
For the smooth channel and Re ¼ 30; 000, the distribution of the local Nu=Nuà in
the vicinity of the turn, is reported in Fig. 9a. Air enters the channel from the lower
duct and leaves from the upper one. By moving streamwise along the channel, the
ratio Nu=Nuà increases around the turn and downstream of it because of the
presence of secondary flows. Three relatively high heat transfer regions may be
recognised: the first one is located by the end wall (in front of the partition wall
towards the first outer corner) and is caused by the jet coming from the first duct
which impinges on this wall; the second one is located at the outer wall downstream
of the second corner and is due to the jet effect of the flow through the bend; the
third one is located at about the half part of the partition wall, downstream of the
second inner corner, where the flow rebounding from the outer wall, impinges before
exhausting. The second zone attains Nu=Nuà values much greater than the other two
due to the presence of strong secondary flows already found by Arts et al. [32]. Two
relatively low heat transfer zones are also observed, one just before the first corner of
the 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 presence
of ribs is clearly evident from Fig. 9b and c, where Nu=Nuà are shown for the two
tested rib-pitch to rib-side ratios P/e. In all the Nu maps ribs are clearly visible due to
the higher heat transfer rate that occurs on them. For both dimensionless pitches, the
thermal pattern before the turn appears to be repetitive (in a sense, the flow could be
considered as thermally fully developed ). For example, in Fig. 9c, the shape and levels
of 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 the
duct entrance.
The secondary flows induced by the V-shaped ribs have the form of two pair of counter rotating cells and produce variations in the spanwise Nusselt number
distribution both in the inlet and in the outlet channel by decreasing the convective
heat transfer coefficient towards the channel axis with respect to that nearby the
walls.
Especially in the inlet duct, the reattachment line downstream of the ribs can be
identified by the locus of the normalised Nusselt number local maxima when moving
in the mean streamwise direction. The reattachment distance, which increases for
the higher rib pitch, appears also to increase going from the walls towards the
channel axis and this is most likely due to the interaction of the main flow with the
secondary one.In the proximity of the first external corner, it is possible to see a low heat transfer
zone, due to a recirculation bubble as already observed for the smooth channel. Just
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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 to
shift downstream for increasing pitch.
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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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For both pitches, the overall increase of the turbulence due to the bend induces
higher values of the normalised Nusselt number in the outlet duct but the percentage
increase 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 problem
of viscous inviscid flow interactions and, in particular, on shock-wave/boundary-
layer interaction phenomena that are of great practical importance for air-breathing
engine inlets, wing/body junctures and deflected control surfaces. Prediction of
thermal and dynamic loads on surfaces exposed to hypersonic flows is an essential
prerequisite for the effective design of aerodynamic control surfaces and of thermal
protection system of modern space vehicles in their trans-atmospheric flight portion.
Measurements presented in this section refer to shock wave-boundary layer
interaction in a two-dimensional hypersonic flow over a model consisting of a flat
plate followed by a compression ramp (wedge) with its hinge line parallel to the
model’s leading edge.
The model surface is realised with two separate MACORTM plates screwed onto
aluminium supports. The model spanwise dimension is 100 mm. The hinge line is
positioned at 50 mm from the leading edge and the ramp angle is 151. MACORTM
was chosen as the model surface material for its low thermal conductivity, as
required 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 a
specific total enthalpy up to 2.5 MJ/kg on an effective test section 60 mm in diameter,
in the low to medium Reynolds number range (104 –106). The tunnel operates in a
pulsed, quasi-steady mode, with running time ranging from 50 to 200 ms. HEAT
facility mainly consists of an arc gas heater and a contoured expansion nozzle,
installed in a vacuum chamber volume of 4.1 m3; auxiliary systems are fitted to the
arc heater to provide it with working fluid and energy. Four rotary pumps evacuate
the chamber until an ultimate pressure of 10 Pa is reached before each run. This
vacuum level allows an under-expanded hypersonic flow-field to be maintained at the
nozzle exit for a running time longer than 200 ms. IR camera used during the test wasFLIR SC 3000 and acquisition frame frequencies was 60 Hz for flow visualisation
and 300 Hz for heat-flux measurements.
A thermal map recorded about 80 ms after the starting of wind tunnel is reported
in Fig. 10. The temperature distribution is almost bidimensional only near the model
leading edge (the flow comes from left to right). Moving downwind, the continuous
decrease of wall temperature, indicates the development of the boundary layer. Near
the hinge line is clearly visible a region where the temperature attains a minimum
that is due to the presence of a separation region in the flow. Moving along the
symmetry axis after the hinge line the temperature reaches a maximum that is to be
correlated to the flow reattachment on the ramp.If one considers that, in a first approximation, the potential core may be
assimilated to a cone emerging from the nozzle exit (the cone height being
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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 thermal
map. 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 alternative
approach proposed in [23] because during the first 30 ms 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 by
means of the Stanton number based on the adiabatic wall temperature computed by
means of the recovery factor for laminar boundary layer flow [23].
Experimental data are also compared with the classical flat plate boundary layer
analytical solution [35]. The results show a good agreement with theoretical solutionon the flat plate. Near the hinge line (X ¼ 50 mm) the presence of a separation region
is clearly identified from the minimum of the Stanton number distribution. The
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Fig. 10. Temperature map (in 1C) recorded on model surface after 80 ms from tunnel starting. Total
enthalpy: 2.3 MJ/kg; stagnation pressure: 4.6 bar.
Fig. 11. Stanton number profile on symmetry axis. Total enthalpy: 1.8 MJ/kg; stagnation pressure: 6 bar.
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entity of the heat flux at reattachment is in good agreement with data present in
literature.
6. Conclusions
The application of infrared thermography as an optical method in heat transfer
and fluid flow visualisation is analysed. The heat-flux sensors, which are normally
used for the measurements of convective heat transfer coefficients, and the
application of the infrared scanning radiometer as a temperature measuring device
are critically reviewed. In particular, the corrections of the errors associated with
tangential conduction along the sensor are investigated for the heated-thin-foil , the
thin-film and the wall calorimeter sensors.The heated-thin-foil heat flux sensor coupled with measurement of surface
temperature by IR thermography is used to measure the convective heat transfer
coefficient on two flowfields: a circular cylinder at an angle of attack and a 1801 turn
channel with and without rib turbulators. Furthermore the thin-film sensor has been
applied to the study of the shock wave/boundary layer interaction in a flat plate with
a ramp in a high enthalpy hypersonic wind tunnel.
For the circular cylinder, in order to measure temperatures, in the whole heated
zone and to account for the directional emissivity coefficient, three thermal images in
the azimuthal direction are taken and patched up. For the sharp edge cylinder and
an 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 Nusselt
number assumes also higher values. On the contrary, for the round nosed cylinder
two thermal reattachment points are present on the leeward side, while no
reattachment is evident at the windward one.
In the inlet zone, ribbed channels show spanwise variations of the heat transfer
maps because of the presence of secondary flows. For both tested rib pitches, the
overall increase of turbulence due to the bend induces higher values of the
normalised Nusselt number, but, in the outlet duct, the percentage increase is lower
than 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 hypersonic
flows has been studied by means of heat-flux measurement performed by IR
thermography coupled with thin-film sensor. The use of IR thermography
demonstrate that the flow condition are two dimensional only geometrically.
However, on the symmetry axis the IR quantitative measurements are in good
agreement 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) 261–281 279
8/3/2019 Basic Thermography
http://slidepdf.com/reader/full/basic-thermography 20/21
[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. NumerHeat 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 didistribuzioni 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) 261–281280
8/3/2019 Basic Thermography
http://slidepdf.com/reader/full/basic-thermography 21/21
[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.
ARTICLE IN PRESS
T. Astarita et al. / Optics and Lasers in Engineering 44 (2006) 261–281 281