An experimental assessment of resistance reduction and wake modification of a KVLCC model by using outer-layer vertical blades

International Journal of Naval Architecture and Ocean Engineering.
2014.
Mar,
6(1):
151-161

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- Published : March 31, 2014

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In this study, an experimental investigation has been made of the applicability of outer-layer vertical blades to real ship model. After first devised by Hutchins and Choi (2003), the outer-layer vertical blades demonstrated its effectiveness in reducing total drag of flat plate (
Park et al., 2011
) with maximum drag reduction of 9.6%. With a view to assessing the effect in the flow around a ship, the arrays of outer-layer vertical blades have been installed onto the side bottom and flat bottom of a 300k KVLCC model. A series of towing tank test has been carried out to investigate resistance (C
_{TM}
) reduction efficiency and improvement of stern wake distribution with varying geometric parameters of the blades array. The installation of vertical blades led to the C
_{TM}
reduction of 2.15~2.76% near the service speed. The nominal wake fraction was affected marginally by the blades array and the axial velocity distribution tended to be more uniform by the blades array.
Previous result of outer-layer vertical blades, effect of blade height on the local skin friction coefficient c_{f} reduction (Hutchins, 2003 ).
Previous result of outer-layer vertical blades, skin-frictional drag coefficient versus Reynolds number for flat plate with blades array (Park et al., 2011 ).
The present study is an extension of the previous research of
Park et al. (2011)
with a view to assessing the feasibility of the outer-layer vertical blades for three-dimensional flow around a ship model. The blades arrays were attached on the flat surfaces of a KVLCC model and a series of model tests in towing tank were carried out to investigate the resistance (C
_{TM}
) reduction efficiency and the improvement of stern wake distribution with varying geometric parameters.
k
KVLCC was selected because plenty of data are available in the open literatures for this type. A KVLCC model with the scale ratio λ = 100 was manufactured out of wood with the principal particulars being listed in
Table 1
. As shown in
Fig. 3
, three rectangular slots were cut out at port, starboard and bottom sides of the hull and then covered with 10
mm
thickness acrylic panels. On each panel, 3
mm
wide groove was machined to facilitate the installation of blades array. The longitudinal location of the blades array (X
_{b}
) was selected approximately at the midsection,
i.e.
, the 10
^{th}
station. This was made in order for the blades to be aligned to the incoming turbulent boundary layer on the midship parallel part of the hull. The spans of the arrays were set as long as possible within the widths of the acrylic panels. The port and starboard arrays were kept underwater with sufficiently small angle of attack to the incoming boundary layer flows in all test conditions.
Principal particulars of 300k KVLCC.
Blades array installation system in the KVLCC model and geometry parameters of outer-layer blades array.
The vertical blades were constructed from 0.2
mm
thin stainless steel plate. The thickness of blades elements amounts to about 5 wall units (t
^{+}
= 5) or 0.01δ. All vertical blades fit into the grooves of the socket, which were cut elaborately by water jet with 4
mm
interval.
Fig. 3
illustrates the geometric parameters of the outer-layer vertical blades, height h, spanwise spacing z and chord length c. In the previous research of
Park et al. (2011)
, the optimal height h was found to scale with the outer variables such as the boundary layer thickness δ. In this study, the blade height h was varied as 4, 8, 12, 16 and 20
mm
while the spanwise spacing z had the values of 12, 16 and 20mm. In designing those geometric parameters, it was required to estimate the boundary layer parameters at the location of the blades array. This was done by employing the empirical correlation of
Schlichting and Gersten (2000)
as follows;
Here, Re
_{x0}
corresponds to the Reynolds number of the fictitious origin and the value Re
_{x0}
= 1.5×10
^{5}
was used. The value of the streamwise distance x is given as X
_{b}
= 1.6
m
, which is the distance between turbulence promoting studs near FP and the leading edge of blades. In addition, the local skin friction coefficient and consequent calculation of the friction velocity u
_{τ}
for the inner scaling was based on the correlation c
_{f}
(x) = 0.455 / {ln (0.06 Re
_{x}
)}
^{2}
given by
White (1991)
.
Table 2
presents flow conditions calculated at speeds near the design speed of 15.5
knots
.
Flow condition in the present study.
Table 3 (a)
and
(b)
represent the nondimensional blade height with respect to the varying freestream velocity based on the outer scaling h/δ and the inner scaling u
_{τ}
h/ν, respectively. Similar data are tabulated in
Table 4 (a)
and
(b)
regarding the nondimensional blade spacing, z/δ and u
_{τ}
z /ν. The blade height was chosen to be smaller than the boundary layer thickness δ . According to
Park et al. (2011)
, blades protruding out of the boundary layer contribute to the device drag only without any skin friction reduction effect. In this study, a total of 15 combinations resulting from 5 heights and 3 spacings were selected for towing tank. Each array is referred to by the blade height and the spanwise spacing,
i.e.
, h16z12 corresponds to the array with the blade height and the spacing being 16
mm
and 12
mm
, respectively. In
Fig. 4
, five blades arrays (h4z16, h8z16, h12z16, h16z16 and h20z16) are shown.
Nondimensional blade height with respect to varying freestream velocity using.
Nondimensional blade spacing with respect to varying freestream velocity using.
Photographs of outer-layer blades arrays used in the present study.
m
in length, 8
m
in width and 3.5
m
in depth. The resistance of KVLCC model was measured at nine towing speeds from 11
knots
to 16.5
knots
. While the speed interval was set at 1.0
knot
from 11 to 14
knots
, it was set as 0.5
knot
from 14.5 to 16.5
knots
. Throughout the experiment, care was exercised to minimize the measurement uncertainty by carrying out drag measurements four times for each test condition. In addition to the resistance measurement, nominal stern wake distribution was investigated at the propeller plane by a 5-hole pitot-tubes rake. The diameter of the propeller plane was 0.1
m
(1/100 scale). Three dimensional wake velocity components were measured at the positions of 0.3
R
, 0.5
R
, 0.7
R
, 0.9
R
and 1.1
R
in the propeller plane.
Fig. 5
indicates the 5-hole pitot tubes rake and differential pressure gauges used in this study.
Photographs of wake measurement setup.
_{TM}
of the baseline case, bare hull without the vertical array in the present study is plotted as a function of model speed V
_{M}
(
m/s
). Uncertainty estimates for the total drag measurements was made by the standard procedure of
Moffat (1982)
. The 95% precision confidence limits for mean statistic were calculated by Student’s t value for four repeatability tests (2 degrees of freedom). Total uncertainty of this system mainly consisted of the uncertainties from the force gages (max ± 0.25%), the water temperature and the towing carriage speed (max ±1%). The measurement uncertainty in the R
_{TM}
ranged ±1.23 to ±0.76%, as depicted in
Fig. 6
.
Total resistance of bare hull without vertical array.
Fig. 7
shows the total drag coefficient C
_{TM}
of total 15 cases with varying blades configuration along with the bare hull case. It is seen that the installation of the vertical array led predominantly to the reduction of C
_{TM}
with some exceptions at lower speeds. Since the vertical blades array is located below the waterline, the variation of resistance is caused by the change of turbulent boundary layer around the ship hull. Therefore, the C
_{TM}
reduction in the present study results from the skin frictional drag reduction by vertical blades array identified by
Park et al. (2011)
. The positive and negative figure at each ship speed represent maximum percentage of C
_{TM}
increase and decrease, respectively. A closer inspection of
Fig. 7
indicates that maximum reduction of C
_{TM}
becomes reasonably uniform for speeds V
_{M}
over 0.66
m/s
(corresponding to V
_{S}
> 13
knots
). Near the design speed, between 0.7 7
m/s
and 0.82
m/s
(corresponding to V
_{S}
= 15~16
knots
in full scale), the blades array led to maximum C
_{TM}
reduction of 2.15~2.76%. It is found that difference in blades geometry causes variation of resistance performance. For instance, the array h16z20 (16
mm
in height and 20
mm
in spanwise spacing) led to resistance reduction of 2.28%, whilst the array h12z12 gave resistance increase of 0.16% at design speed.
C_{TM} versus Reynolds number with various blade arrays.
In order to delineate the effect of blades geometry on C
_{TM}
, the percentage C
_{TM}
reduction (1 – C
_{TM}
/ C
_{TM0}
)×100 is depicted in
Fig. 8
in a parameter space of h/δ, z/δ and Reynolds number of model Rn
_{M}
. The C
_{TM}
of bare hull is noted as C
_{TM0}
. The positive values correspond to C
_{TM}
reduction. The geometric parameters are nondimensionalized using outer scaling method. This is because the optimal height h was found to scale with the outer variables such as the boundary layer thickness δ (
Park et al., 2011
). As seen, three-dimensional parameter spaces are divided to subplanes where the 2-D contour plots are drawn with one parameter being fixed. At a first glance, the C
_{TM}
reduction effect is most noticeable for the lowest speed near 0.57
m/s
(corresponding to V
_{S}
= 11
knots
). At higher speeds, local maxima of the contours are found, implying that there is a certain relationship for an optimal geometry of blades array. A closer inspection of
Fig. 8
reveals that h/δ and z/δ at local maxima of C
_{TM}
reduction is linearly correlated with each other. For instance, h/δ = z/δ = 0.6 and 0.8 show local maxima of C
_{TM}
reduction for V
_{M}
= 0.64 and 0.72
m/s
(V
_{S}
= 12.5 and 14
knots
). The dependence of C
_{TM}
reduction on nondimensional blade height h/δ is given in
Fig. 9
at fixed z/δ. It is notable that the same h and z values give different nondimensional parameters h/δ and z/δ according to the variation of model speed. In
Fig. 9
, the considerable scatter between curves makes it difficult to extract a decisive tendency of C
_{TM}
reduction on h/δ. However, it is observed that the C
_{TM}
reduction tend to decline after a certain value of h/δ. For z = 12
mm
(z/δ ~ 0.6) in
Fig. 9(a)
, the peak of C
_{TM}
reduction is found near h/δ = 0.4~0.6, while the corresponding optimal h/δ is located near h/δ = 0.8 for z = 16
mm
(z/δ ~ 0.8) in
Fig. 9 (b)
. On the other hand, for z = 20
mm
(z/δ ~ 1.0) in
Fig. 9 (c)
, the C
_{TM}
reduction drops after h/δ = 0.8. This supports the linear correlation between optimal h/δ and z/δ up to a certain threshold value, which is also consistent with the flat plate skin frictional drag measurement result (
Park et al., 2011
). As previously addressed by
Park et al. (2011)
, this proportional relationship is based on the balance between mutually competing mechanisms of the attenuaiton of turbulence and the augmentation of device drag.
Contour plots of percentage C_{TM} reduction in nondimensional geometry parameter space; (a) Plots with constant blade height, (b) Plots with constant model speed V_{M}.
Effect of blade height on percentage C_{TM} reduction (a) z = 12mm , (b) z = 16mm , (c) z = 20mm .
As mentioned in the introduction, the local skin friction reduction by means of vertical blades array was first demonstrated by
Hutchins and Choi (2002)
and
Hutchins (2003)
. According to
Fig. 1
showing their results, the local skin friction reduction effect persists up to x
^{+}
≈ 5,000 (x/δ ≈ 12). It is also seen that the maximum local skin friction reduction becomes saturated for blade height h/δ > 0.5. However, from the aspect of downstream persistence, it is implied that the higher blade height, the longer is the skin friction reduction sustained. There are two factors affecting the resistance reduction by the vertical blades in the present study. First, the total drag reduction is an integrated effect of local skin friction reduction along the downstream distance. Second, it is notable that the distance between the vertical blades array to stern corresponds to approximately 80δ in the present study. These factors account for the optimal height h/δ = 0.6 ~ 0.8 observed in this study.
_{n}
in both cases are shown to decrease, but only slightly. The wake distributions also remain essentially the same with slightly noticeable acceleration of axial velocity component V
_{x}
(plotted in colored contours) in the dead water region near θ = 0°. Taking the distance between the blades array and the propeller plane, the marginal effect of blades on nominal wake is quite natural. With a view to investigating the effect of blades array on the wake flow, the axial velocity distributions along the azimuth angle at r = 0.7
R
for all blade cases are plotted against that for the bare hull case in
Fig. 11
. For all cases, it is discernible that the axial flow is accelerated in the range of 0° ≤ θ ≤ 30°, whereas the deceleration takes place for 60° ≤ θ ≤ 160°. In other words, the wake flow becomes more uniform with the presence of blades array.
Comparison of wake distribution for representative blade configurations.
Axial velocity distributions at 0.7R for each spanwise blade spacing (a) z = 12mm , (b) z = 16mm , (c) z = 20mm .
Fig. 12
gives percentage wake fraction reduction (1 – w
_{n}
/w
_{no}
)×100 as a function of nondimensional blade height h/δ for fixed nondimensional blade spacing z/δ. Here, w
_{no}
refers to the wake fraction for bare hull case with w
_{no}
= 0.5789. No conspicuous trend is found, which implies that the influence of vertical blades is not dominant enough to make a significant change in wake flow. In the overall, the effect of vertical blades on wake distribution is only marginal. Installing the blades array nearer to the propeller would lead to substantial change in wake, but that would in turn diminish the resistance reduction effect because the area of downstream influenced area becomes smaller.
Percentage wake fraction reduction against nondimensional blade height.
k
KVLCC model ship. The present study is a sequel to the flat plate skin frictional drag reduction research using the vertical blades array by
Park et al. (2011)
. The blade geometry was designed to follow the outer scaling based on the boundary layer thickness δ, which was found to be effective by
Park et al. (2011)
. The installation of vertical blades led to the resistance (C
_{TM}
) reduction of 2.15~2.76% near the service speed, while the resistance reduction was found to be 2.15~3.73% in the entire range of ship speed under consideration. The linear correlation between optimal blade height h/δ and spanwise spacing z/δ is corroborated again in the model ship experiment. The resistance reduction effect becomes saturated as the nondimensional blade height h/δ approaches 1, which is consistent with the tendency in the previous research (
Park et al., 2011
). The nominal wake fraction was affected marginally by the blades array and the axial velocity distribution became more uniform by the blades array. The minor change of the wake due to the blades array is accounted for the limited downstream influence. In the overall, the drag reduction capability of the outerlayer vertical blades has been manifested by the present model ship experiment. It is remarkable that the present outer-layer vertical blades, unlike most existing hull appendages, could contribute to the improvement of energy efficiency through the resistance reduction.
It is worthwhile to mention that the present vertical blades array has a scaling issue arising from the Reynolds number dependence. Together with the previous study of
Park et al. (2011)
, however, it has been consistently found that the influence of the present outer-layer vertical blades on the boundary layer flow scales with the outer variable,
i.e.
, the boundary layer thickness δ. This strongly implies that the present vertical blades may well be applied to full scale ship to obtain the similar effects as those found in the model study, provided the geometry is designed after the local boundary layer thickness. As addressed in the previous study, the outer scaling device as the present vertical blades array is advantageous in the application point of view. Whilst the inner scale device such as riblets requires sub-millimeter scale which is prohibitively small for real ship application, the present blades array remain in the feasible manufacturing scale of the boundary layer thickness. For instance, the boundary layer thickness at the midship of KVLCC for the design speed of 15.5
knot
is estimated to be 1.3
m
. Therefore, typical optimal blade height of 0.6 boundary layer thickness amounts to 0.79
m
. Considering the streamwise distribution of the local skin friction reduction shown in
Fig. 1
, the skin frictional reduction effect of the vertical blades array persists up to 20 boundary layer thicknesses downstream. Combining δ = 1.3
m
in full scale ship with this leads to the estimates for the effective downstream zone of approximately 10% of ship length in full scale. After this downstream length, the installation of an additional blades array, which needs to be appropriately scaled with the local boundary layer thickness, could lead to an additional skin frictional drag reduction. Although the effect of single blades array in the present study seems to be marginal, it is thus envisaged that the installation of blades arrays at multiple streamwise locations would give rise to more noticeable resistance reduction capability.

Energy efficient ship
;
Skin frictional drag reduction
;
Outer-layer vertical blades
;
Effective power
;
Turbulent boundary layer
;
Wake

INTRODUCTION

Recent increase of oil price has triggered a significant change in the ship design trends, the energy saving design. The efforts to achieve energy efficient ship are divided into two categories; drag reduction and propulsion efficiency enhancement. Historically, more attention has been paid to the hull form optimization toward the reduction of wavemaking resistance than the turbulence control for the frictional drag reduction. However, the frictional drag accounts for about 70~80% of total drag of slow and full hull ships. This makes the frictional drag reduction significant for the energy saving design.
The frictional drag of ship is attributed to the wall shear stress of the turbulent boundary layer flow around the hull surface. The wall shear stress of turbulent boundary layer is in turn closely associated with the organized coherent structures such as hairpin vortices in the turbulent boundary layer flow. Various strategies of turbulent flow control toward either destruction or attenuation of such organized coherent structures have been proposed during several decades. From the viewpoint of reliability, the passive techniques such as riblet (
Walsh, 1980
;
Beckert and Bartenwerfer, 1989
), compliant coating (
Choi et al., 1997
) and Large Eddy BreakUp device (LEBU) is more suitable for the marine application.
Hefner et al. (1979)
conducted the experiments with LEBU to reduce the skin friction downstream of the LEBU devices and achieved 24% of drag reduction compared to undisturbed flat plate levels. The LEBU devices directly interact with and change the large eddy structures, thus interruptting the production loop and reducing the bursting events causing surface stress.
Schoppa and Hussain (1998)
claimed that the imposition of a large-scale vortical control flow into turbulent boundary layer flow could lead to promising skin-friction reductions.
Hutchins and Choi (2002)
and
Hutchins (2003)
introduced the idea of the outer-layer vertical blades, an array of thin vertical plates in the turbulent boundary layer. This array chops off large scale coherent structures of turbulence, thereby disconnecting the link between outer and inner layers. The height and the spanwise spacing were varied to find optimal values. As seen in
Fig. 1
, the skin friction reduction effect attains a maximum of 30% when the height of the blade h is approximately 50% of the local boundary layer thickness δ at the position of the blade in the undisturbed boundary layer without blades (h/δ = 0.5 ). In this case, this maximum occurs at the position of six boundary layer thickness downstream of the blade ( x/δ = 6). Recently
Park et al. (2011)
investigated the frictional drag reduction capability of the outer-layer vertical blades. Whilst
Hutchins (2003)
demonstrated only the reduction local skin friction effect in the downstream of the blades, the towing tank measurements of
Park et al. (2011)
clarified the practical applicability by considering the device drag of the blades. It was found that the outer-layer vertical blades array achieved maximum 9.6% drag reduction, as presented in
Fig. 2
.
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EXPERIMENTAL METHODS

- Model ship and flow conditions

For the present model test, 300
Principal particulars of 300kKVLCC.

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Flow condition in the present study.

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Nondimensional blade height with respect to varying freestream velocity using.

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Nondimensional blade spacing with respect to varying freestream velocity using.

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- Towing tank experiment

The measurements were performed in the towing tank facility of Pusan National University. The dimension of the towing tank is 100
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RESULTS AND DISCUSSIONS

- Effect of geometric variation of vertical blades on the resistance reduction in model scale

In
Fig. 6
, the resistance R
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- Effect of geometric variation of vertical blades on the wake distribution

In the present study, assessment has been extended to the wake distribution affected by the vertical blades array. This is based on the assumption that the change in the turbulent boundary layer which led to the drag reduction might as well affect the wake distribution.
Fig. 10
illustrates such variations in wake distributions for a couple of blade configurations, h4z20 and h20z20, in comparison with the wake in bare hull case. The nominal wake fractions w
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CONCLUSIONS

In the present study, an experimental assessment in towing tank has been made of the resistance reduction and wake modification capability of the outer-layer vertical blades, which is attached to the mid-section of a 300
Acknowledgements

This work was supported by the National Research Foundation of Korea (NRF) grant funded by the Korea government (MSIP) through GCRC-SOP (No. 2011-0030013).

Hefner J.N.
,
Weinstein L.M.
,
Bushnell D.M.
1979
Turbulent drag reduction using compliant surfaces
Viscous flow drag reduction; Symposium
Dallas, Texas
7-8 November 1979
110 -
127

Hutchins N.
2003
An investigation of larger-scale coherent structures in fully developed turbulent boundary layers
University of Nottingham

Hutchins N.
,
Choi K.S.
2002
Towards a greater understanding of turbulent skin friction reduction
Proceedings of ASME FEDSM’02 2002 ASME Fluids Engineering Division Summer Meeting
Montreal, Quebec, Canada
14-18 July 2002
1449 -
1454

Schlichting H.
,
Gersten K.
2000
Boundary layer theory
8th Ed.
Springer-Verlag
Berlin

Walsh M.
1980
Drag characteristics of V-groove and transverse curvature riblets
Symposium on Viscous flow drag reduction
Dallas, TX
7-8 November 1979
168 -
184

White F.M.
1991
Viscous fluid flow
2nd Ed.
McGraw Hill
New York

Citing 'An experimental assessment of resistance reduction and wake modification of a KVLCC model by using outer-layer vertical blades
'

@article{ E1JSE6_2014_v6n1_151}
,title={An experimental assessment of resistance reduction and wake modification of a KVLCC model by using outer-layer vertical blades}
,volume={1}
, url={http://dx.doi.org/10.2478/IJNAOE-2013-0169}, DOI={10.2478/IJNAOE-2013-0169}
, number= {1}
, journal={International Journal of Naval Architecture and Ocean Engineering}
, publisher={The Society of Naval Architects of Korea}
, author={An, Nam Hyun
and
Ryu, Sang Hoon
and
Chun, Ho Hwan
and
Lee, Inwon}
, year={2014}
, month={Mar}