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DOI: 10.1177/0954407011404763

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D M Heim and J B GhandhiInvestigation of swirl meter performance

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Investigation of swirl meter performanceD M Heim and J B Ghandhi*

Engine Research Center, University of Wisconsin-Madison, Wisconsin, USA

The manuscript was received on 17 December 2010 and was accepted after revision for publication on 3 March 2011.

DOI: 10.1177/0954407011404763

Abstract: The performance of vane- and impulse-type swirl meters was investigated, and adirect calibration method for swirl meters was developed. The zero-swirl bias of the meterswas tested by installing an axially aligned tube on the swirl meter. Both the vane- andimpulse-type meters showed insignificant zero-swirl bias. A known swirl was provided to theswirl meters using an offset, inclined tube arrangement. The angular momentum flux deliv-ered by this system was found to depend linearly on the product of the offset distance andcosine of the inclination angle. Both the impulse- and vane-type meters were found to givemeasurements below the known swirl value, but both meters gave results that were linearlydependent on the angular momentum flux, which allows characterization of the meter’s effi-ciency with a single parameter. The efficiency of the impulse-type meter varied from 0.7 to0.93, was a moderate function of the flow straightener aspect ratio, and depended slightly onthe meter size. The vane-type meter’s efficiency was 0.32–0.45 for the conditions tested, wasinsensitive to the paddle wheel flow straightener aspect ratio, and depended significantly onthe meter size. The vane-type meter measurements were also found to depend on the paddle-to-bore-diameter ratio; values slightly exceeding unity should be used. The swirl meter effi-ciency can be used to correct measurements to an absolute basis. Based on these findings, auniversal correction factor does not exist, and a given measuring device will need to be cali-brated using the methodology described.

Keywords: swirl characterization, in-cylinder engine flow, port flow

1 INTRODUCTION

The combustion rate in an internal combustion

engine has long been understood to depend on the

in-cylinder mixture turbulence and the turbulence is,

in turn, directly influenced by large-scale flow struc-

tures such as swirl and tumble. It has been shown

that higher levels of swirl produce higher levels of

turbulence and lower cyclic variation [1]. Thus, the

level of swirl produced in an engine can directly

affect engine performance, cycle-to-cycle variability,

and emissions. Therefore, being able to measure

accurately and easily the swirl characteristics pro-

duced by an engine head is of great importance.

As early as 1934, Alcock [2] described using an in-

cylinder rotating vane to measure an optimal swirl

ratio that gave the best performance for a given

engine. These studies, however, required special

engine heads to accommodate placement of the vane

inside the cylinder and eliminate obstructions due to

poppet valves and injectors. Steady flow tests with a

vane-type meter subsequently replaced the in-cylin-

der rotating vane measurements, and have been used

for decades. Fitzgeorge and Allison [3] measured

swirl speed using a two-bladed impeller inside a flow

rig cylinder. They adjusted the axial distance between

the impeller and engine head and found the impeller

speed was a maximum when this distance was

1.4 times the cylinder bore diameter. They also used

the steady swirl results to try to predict the swirl in

an actual engine. Jones [4] measured swirl speed

using a straight-bladed anemometer inside the

flow rig cylinder and Watts and Scott [5] used a

*Corresponding author: Engine Research Center, University of

Wisconsin-Madison, 1500 W. Engineering Dr., Madison, WI

53706, USA.

email: [email protected]

1067

Proc. IMechE Vol. 225 Part D: J. Automobile Engineering



rectangular-shaped vane in their flow rig cylinder

and noted the form of the vane had little influence

on the measured swirl. Tindal and Williams [6] stud-

ied the air flow patterns in a steady-flow rig using

light paper flags and a vane anemometer to measure

swirl speed. They simulated the presence of a piston

in the cylinder by inserting a restrictor plate into the

flow rig liner at two bore diameters away from the

cylinder head and found that it caused the axial velo-

city to assume a more regular pattern, which resulted

in an increase in measured swirl.

Tippelmann [7] set forth the idea of using an

impulse-type swirl meter with a flow straightener

that converted the angular momentum into a mea-

surable torque. Uzkan et al. [8] described their

impulse-type meter having a honeycomb with small

cells and large aspect ratio capable of straightening

the swirling flow completely. They also note that the

honeycomb should not be inserted into the rig

cylinder (as in references [7] and [9]), but should lie

below it with a larger diameter to eliminate air

blow-by. Swirl measurements were made using dif-

ferent head-to-honeycomb distances. A monotonic

decrease in measured torque with increasing dis-

tance was observed and attributed to cylinder wall

friction. They estimated the rate at which the angu-

lar momentum decays is on the order of 10 per cent

per cylinder diameter of axial distance.

A number of studies have made comparisons

between vane-type and impulse-type meters and in

general conclude that vane-type meters provide

lower swirl coefficients than impulse-type meters.

Tippelmann [7] showed that the readings from a

vane-type anemometer were too small and varied in

magnitude compared with the impulse-type meter.

Monaghan and Pettifer [10] calculated swirl ratios

for four different types of ports using both vane-type

and impulse-type meters. Swirl ratios using the

impulse-type meter were generally 30 per cent great-

er than those using the vane-type meter. Stone and

Ladommatos [11] took cylinder head swirl measure-

ments using both a paddle wheel anemometer

and impulse-type meter and also concluded that the

paddle wheel results fell below those of the

impulse-type torque meter. Snauwaert and Sierens

[12] acquired steady rig swirl measurements with a

paddle wheel anemometer, impulse-type meter and

a laser Doppler velocimeter (LDV) to show that dif-

ferent flow patterns produced over the range of

intake valve lift have varying effects on measure-

ment accuracy. Tanabe et al. [13] tested the same

sized honeycomb using a vane wheel anemometer

and impulse-type meter. The vane wheel anem-

ometer gave swirl numbers below those of the

impulse-type meter, where the level of difference

depended on port type and valve lift. At the maxi-

mum valve lift the swirl numbers from the vane

wheel anemometer all tended to be about 0.8 times

as large as those calculated using the impulse-type

meter.

There have been limited published results on the

effect of flow straightener and paddle wheel geome-

try on measurement accuracy. In one such study,

Tanabe et al. [13] tested honeycomb flow straigh-

teners with cell sizes of 3.2 and 6.4 mm and heights

of 10, 20, and 30 mm. They first measured the drag

coefficients of the flow straighteners with steady

axial flow. Swirl numbers were then measured with

an impulse-type meter for three different cylinder

head port types at maximum valve lift. They found

that the flow straighteners with smaller drag coeffi-

cients tended to measure higher swirl numbers, but

the differences varied with valve lift.

Several investigators have made LDV measure-

ments to compare with steady-flow measurements.

Monaghan and Pettifer [10] took LDV measure-

ments in the steady-flow device to show how both

vane-type and impulse-type meters affect the axial

and radial velocity profiles in the swirl rig. The axial

flow was shown to be highly non-uniform and high-

er towards the outer part of the cylinder. This dis-

credited the assumption of uniform axial velocity

inherent in the use of the vane-type swirl meter cal-

culations, which leads to an underestimation of cal-

culated angular momentum. Kent et al. [14] made

LDV measurements in a motored engine, then inte-

grated the results to find the mean swirl at the end

of induction. The results were approximately 15 per

cent higher than predicted by the impulse-type swirl

meter, but they found their predictions of in-

cylinder swirl based on steady-flow angular momen-

tum flux measurements to be in trendwise agreement

with the LDV measurements in the motored engine.

In the current investigation, two geometrically

identical engine heads have been built to study the

speed- and size-scaling relationships of engine

flows. The length scale ratio between the engine

heads of this study is 1.69. Geometrically similar

engine heads should produce similar levels of swirl

when appropriately non-dimensionalized. In order

to span a wide range of in-cylinder conditions, the

heads are fitted with both normal and shrouded

intake valves. The first step in ensuring the flow

similarity of the heads was to perform steady-flow

measurements. These measurements, which span a

wide dynamic range in swirl level, brought to light

several features of steady-flow swirl measurements

that needed to be resolved in order to assess the

flow similarity between the scaled engine heads.

Vane- and impulse-type meters have been tested,

1068 D M Heim and J B Ghandhi




and an absolute calibration methodology has been

established in order to compare the results confi-

dently from the different sized heads.

2 EXPERIMENTAL APPARATUS

Testing was performed using a SuperFlow 600 flow

bench. The flow bench pulls air into the machine

with a prescribed pressure drop across an attached

test section ranging from 0.25 to 12 kPa (1 to 48 in

H2O). The volumetric flowrate was found from the

pressure drop across a calibrated orifice inside the

flow bench using an inclined manometer. The den-

sity of the air was calculated from temperature and

humidity data acquired using a Mannix model J411-

TH digital hygro thermometer, and the barometric

pressure was measured using a Heise model CM dial

pressure gauge. The engine heads were tested at the

industry-standard pressure drop of 7 kPa (28 in H2O).

The steady-flow swirl testing was performed

using a different swirl adapter fixture for each size

of engine head (see Fig. 1 and Table 1). Hereafter,

the two engine heads and associated components

will be referred to as ‘small’ for the smaller engine

and ‘large’ for the larger engine. The swirl adapter

fixtures, which have H/B = 1.5, are installed between

the cylinder head and the swirl meter. The valve lift

was adjusted using a modified micrometer that

mounted to the engine head. Intake horns, with

radii of curvature large enough to minimize the

pressure drop at the inlet to the intake ports, were

connected to the entrance of the intake ports. The

bore diameters of the swirl adapter fixtures were

the same as the engine cylinder bore. Table 1 gives

the relevant dimensions of the engine heads and

swirl adapter fixtures used with both vane-type and

impulse-type swirl meters.

The vane-type swirl meter used for this study was

an Audie Technology paddle-style swirl meter. The

meter featured a honeycomb paddle wheel made of

polycarbonate plastic with tubular cells. The outer

diameter of the paddle featured a smooth, thin poly-

carbonate sheet wrapped around the honeycomb to

form a continuous cylinder-like shape. The swirl

meter provided an electronic output of two pulses

per revolution, which are also used to determine

both the direction of rotation and the rotation rate

with the addition of an HP model 5315A universal

counter; for all testing, data were collected and aver-

aged over a 40 s period to obtain the mean rotation

rate of the paddle. The relevant dimensions of the

paddle meter are provided in Table 1.

Two impulse-type swirl meters were used for this

study. Tests were first conducted using the impulse-

type meter from the study by Bottom [15]. This will

be referred to as the ‘first’ impulse-type meter. In this

meter, a polycarbonate shaft is fixed on one end and

the other end is attached to a honeycomb flow

straightener consisting of an aluminium honeycomb

matrix (see Table 1 for dimensions). The shaft was

instrumented with two Vishay/Micro-Measurements

torsional strain gauges located 180� apart. The shaft

was designed to deform elastically for low angular

momentum flows. An Omega DMD-465 strain-gauge

amplifier provided the excitation voltage for the

strain gauges and a data acquisition system recorded

the instantaneous voltage at a rate of 10 Hz. Data

were collected and averaged over a 40 s period to

obtain the mean voltage (torque).

A second impulse-type meter was used for this

study, henceforth referred to as the ‘second’

Fig. 1 (a) Vane-type swirl meter test set-up; (b) impulse-type torque meter test set-up

Investigation of swirl meter performance 1069




impulse-type meter. In this meter, a Transducer

Techniques RTS-5 torque sensor was secured at the

bottom and, similarly to the first impulse-type meter,

a shaft was attached at one end to the sensor and on

the other end to a honeycomb flow straightener. The

honeycomb was made of the same material and tub-

ular structure as used in the vane-type meter. The

design of the second impulse-type meter allowed dif-

ferent honeycomb flow straighteners to be easily

tested. A Daytronic model 3270 strain gauge condi-

tioner/indicator provided the excitation voltage for

the torque sensor and the same data acquisition sys-

tem recorded the instantaneous voltage at a rate of

10 Hz. Dimensions of the honeycomb flow straigh-

tener will be discussed in a later section.

Both impulse-type meters were calibrated by

applying a set of known torques to the centre of the

honeycomb flow straightener. For each applied tor-

que, a corresponding voltage was recorded. Before

and after each applied torque, the zero-torque

voltage was recorded and averaged. The average

zero-torque voltage was subtracted from the applied-

torque voltage to obtain the voltage difference. The

voltage difference was plotted against the applied

torques to determine a linear calibration curve.

Calibration data were collected for counterclockwise

torques applied to the honeycomb flow straightener.

3 FLOW PARAMETERS

The flow parameters that will be used to character-

ize the engine heads are the flow coefficient, Cf, the

swirl coefficient, Cs, and the swirl ratio, Rs. The flow

and swirl coefficients are measured at discrete valve

lifts over the range of the cam profile, and are

reported as a function of L/D, where L is the valve

lift and D is the valve inner seat diameter. The flow

coefficient is a measure of the actual mass flowrate

to a theoretical mass flowrate and is defined as

Cf =_m

rVBAv(1)

where _m is the air mass flowrate, r is the upstream

air density, AV is the valve inner seat area, and VB is

the Bernoulli velocity given by

VB =

ffiffiffiffiffiffiffiffiffi2DP

r

s(2)

where DP is the pressure drop across the test sec-

tion. The incompressible relation for velocity is suf-

ficient at the 7 kPa pressure drop, i.e. the Mach

number is 0.32, but for higher pressure drops a

compressible form of the velocity should be used.

The swirl coefficient, Cs, is a characteristic non-

dimensional rotation rate and is calculated for vane-

type meters using

Cs =vB

VB(3)

where v is the vane or paddle wheel angular velocity

and B is the cylinder bore. For impulse-type swirl

meters, the swirl coefficient is calculated from

Cs =8T

_mVBB(4)

where T is the torque measured by the meter. The

swirl ratio, Rs, is a convenient single metric that

takes into account the flow and swirl coefficients

over the entire lift profile of the engine. The swirl

ratio is calculated as

Rs =ph2

vBS

4AV

ÐuIVC

uIVO

Cf Csdu

ÐuIVC

uIVO

Cf du

!2 (5)

where hv is the volumetric efficiency, assumed equal

to 1 for all calculations, uIVO and uIVC are the crank

angle (rad), at intake valve open and intake valve

closed, respectively, and S is the engine stroke. A full

derivation of the swirl coefficient and swirl ratio can

be found in Appendix 2.

Table 1 Relevant dimensions of the engine heads,

swirl adapters, vane-type meter, and first

impulse-type meter

Parameter

Dimensions (mm)

Large engine head Small engine head

B 82.0 48.6H 123.0 72.8D 35.0 20.7Lmax 7.9 4.7DP 132.0HP 15.9DI 165.0HI 64.0Paddle honeycomb 3.7

cell diameter, dP

First impulse torque 6.4honeycombcell size





4 INITIAL MEASUREMENTS

Initial swirl measurements on the geometrically

similar heads were performed using the first

impulse-type meter. The study included using both

standard and shrouded intake valves, where an 180�shroud was used to produce higher levels of swirl.

The flowrate data indicated that the mass flowrate

was well scaled between the two heads, i.e. a certain

level of similarity had been achieved.

From equation (4) it can be seen that for a con-

stant pressure drop (VB) the torque will scale as the

swirl coefficient and a characteristic length to the

third power. Based on the 1.69 scale ratio, and

assuming a worst-case scenario of a swirl ratio of 3

for the large head with the shrouded valve and a

swirl ratio of 0.15 for the small head with a stan-

dard valve (a 20:1 ratio of Cs), a 96: 1 ratio of torque

is obtained. Thus, a measurement device with a

very high dynamic range is required to cover the

entire test range of interest. It was desired to have

a precision of 1 per cent (a 100: 1 signal-to-noise

(S/N) ratio) in the measurements. The S/N ratio

was calculated based on the variability in the mea-

sured torque over the 40 s integration period. For

both heads fitted with the shrouded valves, the pre-

cision criterion was achieved by L/D~0.07. For the

unshrouded valve cases, at the maximum valve lift

the criterion was just satisfied for the large head,

but for the small head the peak S/N ratio was 25: 1

over all L/D. These results motivated the investiga-

tion of a vane-type meter because, intrinsically, a

rotation rate is easier to measure with a wide

dynamic range.

Figure 2 shows the swirl coefficients for both

heads with the shrouded valves as a function of L/D

measured with both the first impulse-type meter

and the vane-type meter. This condition was chosen

because of the high S/N achieved with the first

impulse-type meter. The impulse meter results

show a good degree of similarity – the resulting swirl

ratios were 2.65 and 2.75 for the large and small

heads, respectively. In contrast, the vane-type meter

results showed two disturbing features. First, the

measurements for both heads differed from the

impulse meter results. Second, the results for the two

heads differed quite significantly from each other;

the swirl ratio was 0.57 for the small head and 1.13

for the large head. The former problem is an issue of

absolute accuracy, which will be discussed below,

but the latter is an issue of the operation of the vane-

type meter and is discussed here.

Owing to the difference in the diameters of the

two swirl adapter fixtures, it was thought that there

might be a difference in air frictional losses from the

paddle outside the cylinder bore (the same sized

paddle was used for both heads). The portion of the

paddle outside the cylinder would experience air

friction tending to retard the motion of the paddle,

which is consistent with the lower Cs measured for

the small head. In order to test the effect of air fric-

tional losses on the rotational speed of the paddle,

custom paddles were fabricated of the same honey-

comb material and geometry as the original paddle

wheel, but with a smaller paddle diameter, DP. For

both the small and large heads, the ratio of the pad-

dle diameter to the swirl adapter fixture, DP/B, was

set to 1.2. Figure 3 shows the results of the constant

DP/B tests for the same conditions as Fig. 2. It can

be seen that by controlling DP/B the differences

between the two vane-type meter measurements

have been eliminated, and it may be concluded that

self-similarity has been achieved. It is also possible

that the gap between the paddle and bore adapter

affected the friction, but this was not expressly

tested. There are, however, still differences in the

absolute value of swirl coefficient between the

impulse- and vane-type meter measurements.

5 ABSOLUTE CALIBRATION OF SWIRL METERS

The wide dynamic range required for these experi-

ments suggests that more than one swirl meter may

be needed. However, based on the initial measure-

ments it is clear that using a vane-type meter for the

low range and an impulse-type meter for the high

-1.0

-0.8

-0.6

-0.4

-0.2

0.0

0.2

Cs

0.250.200.150.100.050.00L/D

Impulse Meter Vane Meter

open symbol - small headfilled symbol - large head

Fig. 2 Initial measurements of swirl coefficient usingthe first impulse-type meter and the vane-typemeters with the standard rotor. The measure-ments are for the shrouded valves





range is not a good option unless an absolute refer-

ence can be established against which both meters

can be calibrated. There are two aspects to an abso-

lute calibration, establishing a zero point and deter-

mining the constant of proportionality (assuming a

linear dependence). Additionally, it is useful to moni-

tor the long-term performance of a flow bench, and

the apparatus that has been developed for calibration

can also be used for this.

5.1 Zero-swirl reference

A zero-swirl reference fixture is shown in Fig. 4(a)

with the relevant dimensions given in Table 2. The

zero-swirl reference features a tube that is coaxial

with the swirl adapter fixture and a flat plate that

secures to the top of the swirl adapter fixture. A flow

straightener was installed at the inlet of the tube in

order to help ensure a uniform incoming flow.

Tests were performed at flowrates corresponding

to a pressure drop of 28 in H2O across the test sec-

tion. The data from the impulse meter showed a

small torque offset in comparison to a zero-flow

condition. Converting this into an equivalent swirl

coefficient based on the measured flowrate, the

maximum value was Cs = 0.024, which is small in

comparison with typical values of Cs, even with the

unshrouded valve. The paddle meter results were

more difficult to quantify because the paddle was

essentially stationary, changing position erratically

but not rotating. It is sufficient to say that both swirl

meters were robust relative to a zero-swirl bias.

5.2 Known swirl reference

The second fixture is a known swirl reference and is

shown in Fig. 4(b) with the relevant dimensions giv-

en in Table 2. The known swirl reference features a

tube with its axis offset from the swirl adapter fix-

ture axis and a flat plate that secures to the top of

the swirl adapter fixture. The tube is installed in the

flat plate at an inclination angle uR relative to the

horizontal. Again, a flow straightener was installed

at the inlet of the tube.

For a given geometry (R2 and uR), it can be shown

(see Appendix 3) that the angled-tube geometry pro-

vides a constant value of Cs; the correlation for any

geometry is provided in Appendix 3. Thus, using

equation (4), T can be found, or combining equa-

tions (3) and (4), a measured v can be used to find

an equivalent torque, Teq, as a function of the mea-

sured velocity V, which is determined from the vol-

ume flowrate and pipe area and is used in place of

the Bernoulli velocity. In the subsequent plots, the

term angular momentum flux will be used, which is

equivalent to T (see equation (6) in Appendix 2).

Figure 5 shows the results of the angled-tube cali-

brations of the vane- and impulse-type swirl meters

for both the large and small fixtures (the ‘second’

impulse-type meter was used for these measure-

ments). For the vane meter measurements DP/

B = 1.2 was used, and for all cases the cell aspect

ratio (HI/dI or HP/dP) was 4.3. Both measurement

techniques show excellent linearity with respect to

the angular momentum flux, but there is not a 1: 1

correspondence between the measured (or derived

in the case of the vane meter) torque and the inlet

angular momentum flux. The high degree of linear-

ity indicates that a single conversion efficiency can

be used to describe the performance of the swirl

meters, and this efficiency is the slope of the lines in

Fig. 5. For the data in Fig. 5, the efficiency ranges

from 0.90 for the large fixture using the impulse-

type meter, to 0.32 for the small fixture using the

vane-type meter. From Fig. 5 it is clear that the con-

version efficiency is a function of the meter type

and the fixture size. The impulse-type meter gives

results that are larger in magnitude than the vane-

type meters by nearly a factor of two, and the

impulse-meter results are closer to, but still less

than, the correct value. For the data in Fig. 5, the

smaller fixture gave higher results for both meters.

The effect of the flow straightener or vane cell size

was measured using polycarbonate honeycombs

having a tubular geometry. The honeycomb cell dia-

meters tested were 6.4 and 3.7 mm. For the vane-type

meter DP/B was again set to 1.2 to minimize the fric-

tional losses, and the honeycomb height was limited

-1.0

-0.8

-0.6

-0.4

-0.2

0.0

0.2

Cs

0.250.200.150.100.050.00L/D

Impulse Meter Vane Meter

open symbol - small headfilled symbol - large head

Fig. 3 Swirl coefficient using the paddles with Dp/B = 1.2 for the vane-type swirl meter. Theimpulse-type meter measurements are thesame as Fig. 2





to 15.9 mm (Hp/dp = 4.3) by the meter design. For the

impulse-type meter longer honeycombs were tested,

up to HI/dI = 17, and a fixed straightener diameter of

DI = 104 mm was used. The cell geometry results are

shown in Fig. 6. The vane-type meter (Fig. 6(a)),

showed a weak sensitivity to the cell geometry, but as

was seen in Fig. 5 the conversion efficiency is poor.

For the large fixture, the conversion efficiency was

~0.32 and for the small fixture it was ~0.44. The lower

conversion efficiency for the large fixture could be

due to friction at the hub, which would be greater for

the larger vane size, or from slip between the air and

the paddle. If air slip was causing the low conversion

efficiency, it might be expected that the higher HP/dP

cases would perform better, which was not the case.

The impulse-type meter showed a stronger sensi-

tivity to the flow straightener geometry, with the con-

version efficiency decreasing with increasing aspect

ratio of the honeycomb. This result agrees with the

findings of Tanabe et al. [13]. In comparison with the

vane-type meter, the conversion efficiency of the

impulse-type meter is significantly larger, but differ-

ences exist between the two fixture sizes and the

magnitude of the conversion efficiency can be as low

as 0.7. Thus, the results from an impulse-style meter

Fig. 4 Calibration devices for establishing (a) a zero-swirl reference and (b) a known swirlcondition

Fig. 5 Impulse- and vane-type meter responses to aknown angular momentum flux produced fromthe angled tube for the small and large fixture.For all cases a cell height-to-length ratio of 4.3was used

Table 2 Dimensions of the vertical and angled refer-

ence standards

Parameter Dimensions

uR Vertical reference, 90� Angled reference, 45�SR 127.0 mmDR 19.0 mmLR 445.0 mm





will underpredict the true level of swirl. It is possible

that the losses in the H/B = 1.5 cylinder diameter tube

could account for some of the underprediction seen

with the impulse-type meter.

6 RAW DATA CORRECTION

The swirl meter efficiency, such as those found in

Figs 5 and 6, can be used to correct measurements

to an absolute basis, and for the current study to

remove size-dependent measurement artefacts. The

correction procedure simply involves dividing the

measured torque by the efficiency factor deter-

mined using the angled-tube measurements.

Figures 7(a) and (b) show both uncorrected and

corrected data for the vane- and impulse-type swirl

meters acquired using the large head. The data are

shown using the same axis range to highlight the

dynamic range of the measurements. The vane-type

measurements were made with Hp/dp = 4.3 and the

impulse-meter measurements were made with HI/

dI = 1.4. Similar to the results of Fig. 3, the uncor-

rected vane-type meter results are approximately a

factor of two lower in magnitude for the shrouded

valve case (Fig. 7(a)). After correction, this difference

is significantly reduced, and the vane-type meter

measurements slightly exceed the impulse-type

measurements. The unshrouded valve data, which

exhibit very low values of Cs, are slightly overcor-

rected, but since the swirl level is not very signifi-

cant, this is not too problematic.

7 CONCLUSIONS

A methodology to measure the absolute perfor-

mance of swirl meters was developed. An axial tube

Fig. 6 Swirl conversion efficiency as a function of the cell aspect ratio for (a) the vane-type meterand (b) the impulse-type meter

Fig. 7 Raw and corrected swirl coefficient data for (a) a shrouded valve and (b) an unshroudedvalve and the large head





arrangement was employed to determine the zero-

swirl performance of a meter. Both the vane- and

impulse-type meters tested showed insignificant

zero-swirl bias. An offset, angled-tube arrangement

was developed to measure swirl meter performance

against a known swirl reference. The swirl coeffi-

cient for the angled-tube geometry was found to be

a linear function of the product of the offset dis-

tance and the cosine of the inclination angle.

Impulse-type swirl meters were found to give mea-

sured results closer in magnitude to the known swirl

level than vane-type meters; both meter types

showed a linear dependence on the input angular

momentum flux, allowing calibration using a single

coefficient. The absolute efficiency of the impulse-

type meter was found to be a function of its physical

size, and of the geometry of the flow straightener

used, with lower cell aspect ratios giving higher effi-

ciency. Vane-type meters were found to be sensitive

to the paddle-to-bore-diameter ratio; higher values

of Dp/B give lower measured swirl coefficient due to

excess friction. The efficiency of the vane-type

meter was found to be insensitive to the paddle cell

aspect ratio, but was sensitive to the physical size of

the meter, even with a constant Dp/B.

ACKNOWLEDGEMENTS

Support for this work was provided by the Wisconsin

Small Engine Consortium. The authors’ special

thanks are extended to D. Kilian for designing the

new impulse-type meter and for his help in data

collection.

� Authors 2011

REFERENCES

1 Bracco, F. Structure of flames in premixed-chargeIC engines. Combust. Sci. Technol., 1988, 58, 209–230.

2 Alcock, J. Air swirl in oil engines. Proc. Instn Mech.Engrs, 1934, 128, 123–193.

3 Fitzgeorge, D. and Allison, J. Air swirl in a road-vehicle diesel engine. Proc. Instn Mech. Engrs,1962–1963, 4, 151–177.

4 Jones, P. Induction system development for high-performance direct-injection engines. Proc. InstnMech. Engrs, 1965–1966, 180(Part 3N), 42–52.

5 Watts, R. and Scott, W. Air motion and fuel distri-bution requirements in high-speed direct injectiondiesel engines. Proc. Instn Mech. Engrs, 1969–1970,184(Part 3J), 181–191.

6 Tindal, M. and Williams, T. An investigation ofcylinder gas motion in the direct injection dieselengine. SAE paper 770405, 1977.

7 Tippelmann, G. A new method of investigation ofswirl ports. SAE paper 770404, 1977.

8 Uzkan, T., Borgnakke, C., and Morel, T. Charac-terization of flow produced by a high-swirl inletport. SAE paper 830266, 1983.

9 Davis, G. and Kent, J. Comparison of model calcu-lations and experimental measurements of the bulkcylinder flow processes in a motored PROCOengine. SAE paper 790290, 1979.

10 Monaghan, M. and Pettifer, H. Air motion and itseffect on diesel performance and emissions. SAEpaper 810255, 1981.

11 Stone, C. and Ladommatos, N. The measurementand analysis of swirl in steady flow. SAE paper921642, 1992.

12 Snauwaert, P. and Sierens, R. Experimental studyof the swirl motion in direct injection dieselengines under steady-state flow conditions (byLDA). SAE paper 860026, 1986.

13 Tanabe, S., Iwata, H., and Kashiwada, Y. On char-acteristics of impulse swirl meter. Trans. Jap. Soc.Mech. Engrs, Ser. B, 1994, 60(571), 1054–1060.

14 Kent, J., Haghgooie, M., Mikulec, A., Davis, G.,and Tabaczynski, R. Effects of intake port designand valve lift on in-cylinder flow and burnrate. SAEpaper 872153, 1987.

15 Bottom, K. PIV measurements of in-cylinder flowand correlation with engine performance. PhD The-sis, University of Wisconsin–Madison, Wisconsin,USA, 2003.

APPENDIX 1

Notation

Av valve inner seat area

B swirl adapter fixture bore

Cf flow coefficient

Cs swirl coefficient

dI diameter of impulse torque meter

honeycomb cells

dP diameter of paddle meter honeycomb

cells

D inner seat diameter

DI diameter of impulse torque meter

honeycomb flow rectifier

DP diameter of paddle meter paddle wheel

DR diameter of reference standard tubes

H height of swirl adapter fixture

HI height of impulse torque meter honey-

comb flow rectifier

HP height of paddle meter paddle wheel

L valve lift

Lmax peak valve lift

LR calibration tube length

_m mass flowrate of air

P pressure





Rs swirl ratio

R1 radius of angled reference standard tube

R2 angled reference standard offset from the

centre-line of the cylinder bore

S engine piston stroke

S/N signal-to-noise ratio

SR flow straightener length

T measured torque

Teq equivalent torque

V measured velocity

VB Bernoulli velocity

hv volumetric efficiency

uIVC crank angle at intake valve closed

uIVO crank angle at intake valve open

uR angle of reference tube

r density of air

v paddle wheel angular velocity

APPENDIX 2

Swirl coefficient and swirl ratio determination

Consider a cylinder closed at one end and open at

the other end. Fluid enters the cylinder through

some arbitrary surface S1 on the closed end, and

flows uniformly (in the axial direction) out of the

open end of the cylinder. If the exit plane contains a

flow-straightening device such that the exit flow is

purely in the axial direction, then the torque

required to hold the flow straightener is found from

the conservation of angular momentum as

Tz =

ð ðS1

rvurv � dA(6)

If instead, the flow exits the open end of the cylin-

der with a solid-body rotation at rotational rate v,

then using angular momentum conservation the

rotation rate can be written in terms of Tz from

equation (6) as

v =8Tz

_mB2(7)

The swirl coefficient is defined as v normalized by a

characteristic rotation rate VB/B, which using equa-

tion (7) gives

Cs =8Tz

_mVBB(8)

The swirl ratio, Rs, is found by considering the

unsteady angular momentum conservation for the

cylinder

0 =∂

∂t

ð ð ð8

r3vð Þrd8

24

35+

ð ðS1

rvurv � dA(9)

The second term on the right-hand side is just Tz

using equation (6), and by assuming a quasi-steady

filling process for an initially empty cylinder, equa-

tion (9) may be integrated to find

ð ð ð8

r3vð Þrd8=

ðtIVC

0

Tzdt (10)

Assuming that the cylinder contents have a solid-

body rotation at O at the time of intake valve clo-

sure, and that the engine rotation rate is Oeng, then

the swirl ratio is found as

Rs[O

Oeng=

32

rpSB4O2eng

ðuIVC

uIVO

Tzdu (11)

Substituting for Tz from (8) and collecting in terms

of Cf, the following is found

Rs =4V 2

B Aref

pSB3O2eng

ðuIVC

uIVO

Cf Csdu (12)

Where Aref is the reference area used to define Cf.

In order to remove the engine speed from the

denominator of equation (12), the following is noted

ðuIVC

uIVO

Cf du =Oeng

rVBAref

ðtIVC

0

_mdt (13)

and that the rightmost integral is just the mass

delivered per cycle. Assuming that the swept volume

is close to the total cylinder volume, and using the

volumetric efficiency, hv, the following is found

Oeng =4VBAref

hvpB2S

ðuIVC

uIVO

Cf du (14)

Introducing equation (14) into equation (12), the

following is found

Rs =h2

vpSB

4Aref

Ð uIVC

uIVOCf CsduÐ uIVC

uIVOCf du

� �2 (15)

The results presented herein used the inner seat

area for Aref, but inspection of equation (15) sug-

gests that if Aref were chosen to be the cylinder





cross-sectional area pB2=4 further simplifications

would be achieved.

APPENDIX 3

Known swirl calibration

The torque measured by a flow straightener can be

written as

Tz =

ð ðS1

rvurvndA(16)

where the normal velocity, vn = V cos u, the tangen-

tial velocity, vu = V cos u cos a, a is the angle that

the differential area element dA makes with the ver-

tical in Fig. 4(b), and V is the measured velocity

obtained from the flowrate measurement of the flow

bench and the known pipe area. Normalizing all of

the dimensions by the cylinder radius (B/2) and

denoting dimensionless distances with an overbar,

e.g. �R1 = R1= B=2ð Þ, and recasting equation (16) in

terms of the swirl coefficient, the following is found

Cs =4

p �R21

sin u cos u

ð ðS1

�r cos a d �A (17)

The integral was evaluated numerically, and the

results were found to be independent of tube size

�R1, linearly dependent on tube offset �R2, and depen-

dent on the cos u. The results can be summarized in

the single plot shown in Fig. 8, where a single line

that passes through the origin fits all of the data

with a slope as shown.

These results have been used in conjunction with

equation (8) to find the torque as a function of vol-

ume flowrate, where V is used in place of VB.

Similarly, using equations (7) and (8), it is possible

to find the vane rotation rate as a function of vol-

ume flowrate.

Fig. 8 Swirl coefficient dependence on angled tubegeometry





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