Brake and Engine Cooling Flows: Influences and Interactions

Dr. R H Barnard and Prof. P R Bullen ACM Engineering, University of Hertfordshire Dr. Jun Qiao Jaguar Cars Ltd. SYNOPSIS Aspects of the interaction between engine cooling and brake cooling flows have been investigated. The practical and theoretical implications of attaching the engine cooling outlet flow to the underside of the vehicle have been studied, and show significant advantages both in terms of drag reduction and potential for improved brake cooling flow. Wind tunnel tests on a nominal scale generic vehicle with modelled radiator and internal flow were conducted. It was found that the simple expedient of radiusing the rear edge of the cooling flow outlet aperture was sufficient to promote attachment of the cooling flow outlet jet, and that this reduced the cooling flow drag and the front lift. It also lowered the front wheel well pressure, thus improving the potential brake cooling effectiveness. This finding has important implications, for the design of the cooling system and underbody geometry, which are fairly straightforward to implement. The usefulness and limitations of established theoretical relationships for cooling flow efficiency have been investigated, and relevant comments are included. 1 NOTATION Af model projected frontal area Ac Radiator core area AO Cooling flow outlet aperture area CD drag coefficient CpO static pressure coefficient at outlet kp radiator core pressure loss coefficient pstag free-stream stagnation pressure pstatic static pressure

p free-stream static pressure pO static pressure at outlet Vc radiator normalising or "core" velocity VO outlet air speed V free-stream air speed (synonymous with car speed on road) inclination of outlet jet to horizontal inclination of outlet plane to vertical air density 2 BACKGROUND The provision of adequate brake cooling can be a major problem on high performance cars. For such vehicles it is normal to provide a supply of cooling air through a duct connected to an aperture at the front. Since the flow at intake will have lost little or no energy, the flow rate will be governed by the duct size and loss characteristics, and the static pressure in the wheel well. An understanding of the factors affecting this static pressure is therefore vital for the design of the brake cooling flow system. The impetus for the work described here came primarily from a desire to study the factors influencing brake cooling flows. However, the work also provided an opportunity to look at the effects of promoting attachment of the engine cooling air outlet flow. From theoretical considerations, described later, it was expected that such attachment would considerably reduce the drag increment due to cooling. It was appreciated that with an attached outlet flow, some significant revision of the commonly accepted expressions for cooling drag would be necessary. Previous work by Barnard et al [refs. 1, 2 and 3] had validated existing theories only for cases with separated outlet jets. 3 LIMITATIONS OF THEORETICAL EXPRESSIONS FOR DRAG DUE TO

ENGINE COOLING FLOW Although expressions for cooling drag flow were developed for aeronautical purposes some time ago, the expressions most commonly used in current automotive practice are based on a paper published by Soja and Wiedemann [4]. A modified version of this was presented by Barnard [3 and 5] as

cos cos (cooling)f

OpO

c

O

c

f

ccD A

ACVV

AA

AA

VVC

=

12 (1)

This expression is based on consideration of the momentum changes and pressure differences across a control volume illustrated in figure 1. This version of the expression allows for the fact that the outlet aperture is not generally perpendicular to the outlet jet flow. In practice most of the drag force is attributable to the momentum change. In vehicle arrangement shown in figure 1, the aperture area is horizontal, so = 90o, and the pressure term thus has no influence in the drag.

Figure 1 The control volume used for equation 1, illustrated for the case of the

Ahmed shape model, as described in refs. [1 and 2].

The expression has been shown to work well for special cases (Barnard [1, 2 and 3]), but is based on some assumptions that are not valid in many practical cases. There are two primary defective assumptions. 1. The first is that drawing air through the inlet to produce a cooling flow does not

influence the external flow of the basic vehicle shape. Clearly, there can be strong influences both in the intake and outlet regions. This problem has been addressed by Williams [6 and 7] who has introduced appropriate modifications to the momentum/pressure equation.

2. The second assumption is that the cooling air leaves the control volume in the

form of a jet emanating from the outlet aperture, as illustrated in figure 1. The second assumption was valid for the Ahmed model tested in refs. [1 and 2], since the outlet was very close to the rear of the model. It was also partly true for the case of the sports car tested by Barnard [3], since the outlet aperture was sharp and the flow did not tend to attach to the vehicle body. However, if the flow does tend to attach, it is necessary to use a different control volume, as illustrated in figure 2. The problem with this model is that the outlet jet velocity and area are no longer easy to estimate, as they are not simply related to the geometry of the outlet aperture. Although equation 1 now contains an unknown exit area AO, and an unknown exit velocity VO (see figure 2), it is still useful in that it shows that with attached flow, the outlet jet should be nearly axial ( = 0 in equation 1), so the axial momentum change is at a minimum for a given outlet jet speed (VO). The cooling drag penalty should thus be small.

VO

V VVc

Figure 2. An appropriate control volume for the case of an attached outlet jet. The outlet area Ao is now not generally equal to the area of the engine bay outlet aperture.

The experimental result given here indicate that if the control volume outlet velocity is assumed to be axial in direction and similar in magnitude to the jet speed at the outlet aperture, then equation 1 will produce a useful first estimate of the cooling drag. The justification for adopting these assumptions and the attendant limitations are given later. The tests described here involved a simplified vehicle with a well-rounded and ideally situated intake aperture, so modifications to the external flow at intake were probably small, though no attempt has yet been made to validate this experimentally. 4 OTHER RELEVANT THEORETICAL EXPRESSIONS The flow speed through the radiator is controlled primarily by the conditions at the outlet aperture: namely the outlet pressure coefficient CpO and by the ratio of the radiator core area to the outlet aperture area Ac /AO. The core speed Vc may be obtained from

pO

c

pOc

kAA

CVV

+

=

2

1 (2)

This expression is derived in reference [2]. The radiator pressure loss coefficient kp is the normalised stagnation pressure loss through the core: defined by

= 25.0 ctotal

p Vpk (3)

O Voo AoVo

5 THEORETICAL FACTORS CONTROLLING THE BRAKE COOLING FLOW For a ram-driven cooling air flow, the mass flow will be determined by the area and speed of the jet of cooling air. The maximum theoretical speed that can be obtained depends on the stagnation pressure at intake pstag and the static pressure surrounding the outlet jet pstatic. If the intake is placed on the front surface of the vehicle, the stagnation pressure will be close to that of the free stream. Thus, it will be seen that the controlling influence will be the static pressure in the vicinity of the outlet jet near the brake disc, which is in turn defined by the pressure field in the wheel well. The maximum mass flow rate is given by

( )staticstag ppAVAm == 2 (4) In our preliminary studies, we found that for a conventional front-engined vehicle, the wheel well pressure is significantly influenced by the engine cooling flow. This is hardly surprising in view of the close proximity of the cooling flow outlet to the wheel well. In this investigation, the relationship has been studied in some detail. 6 EXPERIMENTAL ARRANGEMENT Rather than study any particular production vehicle, it was decided that a simplified generic vehicle model would be constructed in roughly scale. The model is shown in figures 3 and 4, and takes the form of an MPV which gives an aerodynamically simple shape.

Figure 3 The scale MPV model shown on the moving belt facility.

Core area AC = 0.0088 m2 Frontal area Af = 0.1273 m2

Tunnel speed V =18.5 m/s

Figure 4. Internal arrangement and relevant dimensions of model The underside was mostly flat, but with a diffuser section starting slightly upstream of the rear wheels. An engine cooling radiator was modelled using the same system of wire mesh and honeycomb described in ref. [1]. The cooling air intake was of a simple rounded rectangular cross-section with smooth entry, and the short intake duct has zero diffuser angle. The outlet aperture was sharp-edged and rectangular, with the area being varied by means of a sliding plate. A rounded roughened fairing could be added to the downstream edge to promote attachment of the outlet jet to the vehicle underside. The wheel wells were modelled with dimensions typical of a production vehicle. The model was suspended over a fixed groundboard and used an internal two-component force balance with an external force balance connected to a rear sting to provide pitching moment information from which the rear-wheel/front-wheel lift distribution could be determined. It is intended that further measurements will be made using the moving belt facility, but for present purposes, the fixed ground was employed, as it provides a simpler and quieter arrangement. The purpose of the work was to provide comparative rather than absolute data. In addition to force measurements, several internal pressure tappings were provided, and seven of these were attached to a bank of transducers. For reasons of hardware/software compatibility, pressure and force data were measured consecutively rather than simultaneously.

static tappings

radiator core

pitot tube

adjustable aperture slide

optional rounded edge

3 5 7

6

Pitot-static tube

4 1&2

Static pressures were measured at the following positions: 1. on the inside face of the front wheel well close to the axle line, 2. at a position close to the wheel hub, corresponding to the location of a brake disc, 3. on the underside just downstream of the cooling air outlet aperture, 4. on the underside upstream of the cooling air outlet aperture, 5. near the bottom of the rear face of the vehicle (to determine the base pressure). 6. in the wall of the inlet duct, 7. in the centre of the inlet duct facing forward as a pitot tube so as to record the

inlet flow stagnation pressure. 7 DETERMINATION OF COOLING AIR MASS FLOW RATE The two latter tapping (6 and 7) allowed estimates of the cooling air mass flow rate to be made by using equation 4 with pstatic now being the static pressure in the intake duct. Dividing this mass flow rate by the duct cross-sectional area and the air density, the mean intake flow speed can also be determined. The radiator core reference velocity VC can then be determined from the ratio of the intake duct area to the core area. This ratio was 1.125 for the first set of tests (results shown in Table 1) and 1.0 for the second set (results shown in Table 2). Knowing VC and the free-stream velocity V (which was measured with a pitot-static tube), the velocity ratio Vc/V may be calculated. 8 THE EFFECT OF ATTACHING THE COOLING FLOW OUTLET JET TO THE

UNDERSIDE For the first set of experiments, the outlet aperture was set with an area exactly equal to the core area. The model was run both with the inlet blanked, and with it open. Two sets of tests were conducted, one with a sharp-edged rectangular outlet (the baseline configuration), and the other with the radiused downstream edge. For the second case, it was found that with an unchanged outlet aperture area, there was an increase in the mass flow rate. The aperture was therefore reduced so as to maintain the same mass flow rate as for the baseline sharp-edged outlet. Table 1 shows the results. CD represents the difference in vehicle drag coefficient between the radiator open and blanked cases, and thus the drag increment due to the effects of the cooling flow. Theoretical values of Vc/V are given, and also the theoretical momentum change component of the drag increment due to cooling, for the cases of an axial outlet jet flow direction ( = 0o) and for a vertical outlet direction ( = 90o). The theoretical values of CD and Vc /V are based on equations 1 and 2. In calculating the theoretical velocity ratio, the measured pressure coefficient just upstream of the outlet aperture (from tapping 4) has been used as the value of CpO. The theoretical drag coefficient increment figures are based on the experimentally measured values of Vc /V. The experimental values of Vc/V are probably too high, as no account has been taken of the intake boundary layer. A discharge coefficient of around 0.98 might be appropriate for the determination of the core velocity. This would bring the theoretical and experimental values more into line.

Table 1. Wind-tunnel and theoretical data for model with AC/AO = 1

Pressure tapping No.

Intake blanked

Sharp edge outlet aperture AC/AO = 1

Rounded rear edge outlet aperture AC/AO = 1

Change in CD between blanked and open radiator CD

0.039

0.032

0.0306

0.0308

Theoretical CD based on measured values = 0o of Vc /V = 90 o 0.0456 0.0462

CL (front wheels) +0.0645 +0.028 CL (rear wheels) -0.0639 -0.0691

2 Cp (wheel disc) -0.268 -0.279 -0.282 1 Cp (wheel well) -0.198 -0.132 -0.142 3 Cp (underside downstream) -0.323 -0.878 -0.810 4 Cp (underside upstream) -0.393 -0.134 -0.190 5 Cp (base) -0.217 -0.224 -0.225 Vc /V (measured) 0.330 0.334 Vc /V (theory:- equation 2) 0.321 0.329 Cooling air mass flow rate

(kg/s) 0.067 0.067

It will be seen that radiusing the outlet aperture to provide well attached, and thus axial outlet flow ( =0 in equation 1), reduces the drag increment due to cooling (CD), and that the increment is now quite close to that predicted by equation 1 for an axial jet. There is no reason why this should necessarily be the case, because the speed and area of the cooling air flow as it leaves the control volume will not generally be the same as that at outlet from the aperture. However, it will be seen that the underside pressure coefficient just upstream of the outlet aperture -0.190 is quite similar to the base pressure coefficient in the region where the flow leaves the control volume -0.225. The control volume outlet and aperture outlet speeds should thus be similar. The attached outlet flow will of course be subjected to viscous drag, but this merely replaces a similar drag on the underbody flow. The base pressure falls slightly when the cooling flow is on, and this will result in a small additional drag increment of approximately 0.001. The influence of the outlet aperture radiusing may be seen in the change in the value of the pressure coefficient just downstream of the aperture. For the sharp-edged case, there is a strong separation bubble, giving rise to a strongly negative coefficient value. With a wool-tuft wand, it was easy to detect the presence of this separation bubble. Partial reattachment was seen to occur, but with a wide unsteady flow. For

the radiused-edge case, a clean attached underside flow was found. The overall lift coefficient is not changed much by the radiusing, but there is a change in the front/rear lift distribution. Both the front and rear wheel lift coefficients are lower for the case of the radiused rear edge, with the greater effect being at the front wheels. This is consistent with the reduction in the underbody static pressure upstream of the aperture that occurred with the radiused edge. Also, due to the outlet jet being near axial, a reduction in the lift is to be expected, since there is no cooling flow vertical momentum at outlet. 9 WHEEL WELL PRESSURES The influence of the cooling flow on the wheel well pressures may be seen Table 1. In general, the presence of the cooling flow serves to raise the wheel well pressure, and hence restrict the potential flow speed for the brake cooling air. However, it will be seen that attaching the cooling air outlet flow, by means of the radius, improves the situation. This is fortunate, as it means that with the radiused outlet, one can have improved brake cooling as well as a reduced cooling drag penalty. 10 EFFECT OF CHANGING THE CORE TO OUTLET AREA RATIO AC/AO The effect of increasing this area ratio from 1 to 2 was investigated both for sharp and radiused edge cases. From equations equation 1 and 2, it is predicted that this will reduce the drag due to cooling. It will also however reduce the core flow velocity, so to compensate, the core area has to be increased to preserve the mass flow rate. In the case of our model, an attempt was made to achieve this by removing a small partial banking plate which had been deliberately introduced for the tests with the smaller area ratio. Removing this increased the core area to 0.0099 m2. The predicted reduction in drag is small if the outlet jet is vertical, since the same mass flow rate is involved, so the momentum changes should be unaltered. If the outlet flow is nearly axial (due to attachment), however, there should be a significant reduction in drag as the outlet jet momentum is increased, and the horizontal momentum loss is thus reduced. Pressure and force data were again collected for both a sharp edged outlet aperture and for a radiused rear edge aperture. The results are shown in Table 2. For the radiused outlet case, the measured value of CD again tends towards that predicted from equation 1 for = 0o, and is usefully lower than for the case where the area ratio Ac /AO =1. The theoretical values are only an approximate guide, for the reasons given earlier. In addition, it is difficult to be certain about what value of CPO to use. For present purposes, the underside pressure just upstream of the outlet aperture has been used. The lift coefficient increments show the same trend as in table 1

Table 2. Wind-tunnel and theoretical data for model with AC/AO = 2

Pressure Port no.

Intake blanked

Sharp edge outlet aperture AC/AO = 2

rounded rear edge outlet aperture AC/AO = 2

Experimental change in CD between blanked and open radiator CD

0.0365

0.027

0.0186

0.0186

Theoretical CD based on measured values = 0o of Vc /V = 90 o 0.0468 0.0467

CL (front wheels) +.064 +.031 CL (rear wheels) -0.052 -0.050

2 Cp (wheel disc) -0.268 -0.177 -0.192 1 Cp (wheel well) -0.198 -0.132 -0.134 3 Cp (underside downstream) -0.323 -0.956 -0.864 4 Cp (underside upstream) -0.393 -0.149 -0.209 5 Vc /V (measured) 0.301 0.3005 Vc /V (theory:- equation 2) 0.287 0.294 Cooling air mass flow rate

(kg/s) 0.0675 0.0675

11 INLET EFFECTS AND OTHER CAVEATS For these tests, the intake was deliberately designed to produce predominantly attached flow for both the blanked and open conditions, and is not representative of the kind of complex intake geometries that result from styling and packaging constraints on real vehicles. The influence of intake geometry on the drag increment has not yet been properly investigated by the authors although some tentative measurements were made that showed that the intake did influence the results. This is to be expected, since the intake flow can affect the attachment of the external flow in appropriate cases. The important influence of the intake effects has been studied by Williams [6 and 7]. Apart from the inlet geometry, the model used was untypical of current real road vehicles, in that it had a largely smooth flat underside. The attachment produced by the radiused outlet would therefore be more effective than on a vehicle with a typically rough and geometrically complex underside. This said, however, a measurable improvement was obtained in a production saloon car fitted with a radiused outlet (not reported, for commercial reasons). It is also likely that there will be increasing attention to smoothing the underside of vehicles in the future, since this is one area where significant aerodynamic improvements can be made and, most importantly, made without upsetting the stylists. Full underfloor fairings are now quite

often fitted to high-performance vehicles. Apart from the aerodynamic advantages, such under-trays help reduce drive-by noise. 12 SUMMARY AND CONCLUSIONS It has been found that the drag due to the cooling flow can be reduced by attaching the outlet flow to the underside of the vehicle. The simple expedient of radiusing the rear edge of the outlet aperture was sufficient to make a worthwhile improvement for the vehicle studied here. This may not work for all underbody geometries. Further drag reductions can be made by increasing the ratio of core area to outlet area ratio Ac /AO. This will only be effective if the outlet flow approaches the axial direction. For a vertical discharge, the cooling drag is constant for a given mass flow rate, since that mass of air always loses its axial momentum. The simple theoretical expression of equation 1 is not really valid for an attached outlet flow, but although it does not allow accurate predictions to be made, it still indicates the parametric trends. Attaching the outlet flow for a front-engined vehicle may lower the wheel-well pressure, and hence improve the potential for brake cooling flow. For this vehicle, attaching the outlet flow reduced the front lift, as expected. This is a beneficial effect, as reducing the front end lift can be quite difficult to achieve without changing the styling. More work is required in order to investigate the effect of the intake geometry, and the influences of underbody roughness. 13 ACKNOWLEDGEMENTS The authors would like to acknowledge Jack Williams for sharing his information and thoughts on the subject, and for his helpful criticisms and suggestions. 14 REFERENCES 1. Barnard R. H., and Ledakis N., Physical modelling and optimisation of radiator cooling flow systems, Procs. 2nd MIRA Conference on Vehicle Aerodynamics, Session 5, October 1998. 2. Barnard R.H., Theoretical and experimental investigation of the aerodynamic drag due to automotive cooling systems, Procs. of the Institute of Mechanical Engineers, Journal of Automobile Engineering, Proceedings Part D, vol. 214, no. D8, 2000, pp. 919-927. ISSN 0954-4070.

3. Barnard R.H., Minimising the cooling system drag for a small sports car, Procs. 3rd MIRA International Vehicle Aerodynamics Conference, (MIRA 2000), 18-19 October 2000, Rugby, U.K. 4. Soja H., and Wiedemann J., The interference between internal and external flow on road vehicles, Ingenieurs d'Automobile, Sept 1987 pp. 101-105. 5. Barnard R. H., Road Vehicle Aerodynamic Design, MechAero Publishing, 2001, ISBN 0-9540734-0-1. 6. Williams, Jack, Walt Oler, and Dinakara Karanth, Cooling Inlet Aerodynamic Performance and System Resistance, SAE report 2002-01-0256, March 2002. 7. Williams, Jack, Aerodynamic drag of engine-cooling airflow with external interference, Procs. 4th MIRA International Vehicle Aerodynamics Conference, (MIRA 2002), 16-17 October 2002, Gaydon, U.K.

Figure 1 The control volume used for equation 1, illustrated for the case of theFigure 2. An appropriate control volume for the case of an attached outlet jet.Figure 4. Internal arrangement and relevant dimensions of model7 DETERMINATION OF COOLING AIR MASS FLOW RATE7. Williams, Jack, Aerodynamic drag of engine-cooling airflow with external interference, Procs. 4th MIRA International Vehicle Aerodynamics Conference, (MIRA 2002), 16-17 October 2002, Gaydon, U.K.