CHINA SHIP SCIENTIFIC RESEARCH CENTER
Preliminary Study on Performance Characteris-tics of Propeller and Fixed Guide Vanes
Zhang Jianhwa, Dong Shitang
September 1985 CSSRC Report
English version-85005
(Extracted and translated from original paper published in Journal of Ships of Jiangsu No.2,19R3, 4 in Chinese)
P. 0
.BOX 116, WUXI, JIANGSU
Contents page
Abstrat i
Introduction i
Theorectical calculation n'ethod of the Propeller arid FGV i
Propulsor models and test 4
Results and analysis 7
Conclusions 10
Acknowledgments 10
Synibols li
Abstract
The results of a preliminary study on the performance of combined pro-pulsor consisting of a conventional propeller and fixed guide vanes placed behind the propeller are presented. Based on the results
of
theoretical calculations, a pair of propulsor models was designed and
tested in the cavitation tunnel of CSSRC. The efficiency was compared with those of Troost B-series propeller and AU-series propeller
designed at the sane operating conditions and the same thrust
require-ment. It shows that installation of fixed
guide vanes has a signifi-caritly effect an recovering energy losses in the slipstream.
Introduction
Installation of fixed guide vanes (hereafter referred to as simply FGV) behind a propeller is one of the measures for recovering part
of
the rotational energy losses in the slipstream due to the tangential induced velocity. It was pointed out in2} that
such kind of
pro-pulsor was studied in the 1930's and no further investigation had been rementioned untill 1966, then Grim's paper 2} was published. In
2}, the calculated efficiency of the combined propulsor has been con-trasted with those of conventional propeller and other propulsors, such as propeller combined with vane wheel. It showed that the
propeller combined with FGV achieves an efficiency 5.32% higher than that of
con-ventional one, while the efficiency of the propeller combined with freely-rotating vane wheel is 4.38% higher. It seems that the FGV is
more effective in recovering energy.
For the purpose of making a preliminary study on this device,
theoret-ical calculation based on lifting-line
theory was performed and the
necessary data of design provided. A pair of propeller models
was
tested in the cavitation tunnel of CSSPC.
The results of tests
indi-cate that this combined propulsor has significant effect in
regainning part of the rotational energy losses in the slipstream and
increasing efficiency. At a certain operating condition, its thrust can be
in-creased by 4.5% and propeller efficiency can be increased by
8.4% com-pared with B-series propeller at the optimum design condition.
Theoretical Calculation Method of the Propeller and FGV
Theoretical calculation method
U
for contra-rotating propeller design is applied in this paper, while a set of FGV is regardedas a non-ro-tating aft propeller. And then the interference
velocities between the forward and aft propeller can be obtained. The geometry of com-bined propulsor can be designed and its performance
can he predicted. Some simplified assumptions as mentioned in
f
i} are made, i.e. only circumferential mean of the interference velocities was taken into
ac-count. Therefore, it is
a function of radial coordinate only, when the distance between two propulsors is specified. Now a closed curve on a surface of cylinder with radius r and a common axis with propel-ler is taken as shown in Fig. i(a).
Fig.1 Closed circuit and control surface
S1 and S2 represent the discs at the propeller and FG1 respectively. The tangential induced velocity should vanish everywhere (namely utO) behind the FGV if the rotational energy losses were completely
regain-ed. It is unlikely to achieve uO because of finite number of baldes
and circumferential variation of
u.
Let y be advanced velocity, n be the rotational speed of the
propel-ler. SA cylindrical control surface with infinite radius is taken as shown in Fig 1(b). Its end surface SA and SB are far ahead and far behind the propulsor respectively. When the propulsor travels a dis-tance v/n during one revolution, the increase in energy of the fluid inside the control surface should he
E.
= p ff1
(u2 + u2 + u2)dT (1)i a t r
T
where T is the volume far behind the propulsor and enclosed by cylin-drical surface, the planes SB and S' . E. is the energy loss during the period. But u may be ignore in ¿e field far behind the pro -pulsor, because
norcontraction
occures there. ThenE1=
pff!
(u2+u)dT
(2)
= p ff1
u2dT + pff1
udT
In principle, a definite distribution of u will be induced once thrust, number of blades, diameter and certain ciru1ation distribution are
specified. Therefore authors try to seek such a circulation distribu-tion of the stator that it makes energy loss due to tangential induced velocity minimum. Namely
p 111
udT = Min.T
with constraint condition of total thrust of combined propulsorconst. This variation problem is too complex to solve rigorously. Authors
in-tend to seek the circulation distribution of the stator (suppose the
-2-circulation distribution of the propeller has been defined) by ari
ap-proximate approach.
At first, the
case of propeller
with
infinite
number of baldes is considered.
In this
case,
u
keeps
constant
along each circumference and
g1, F1
(r) +
g2cG F2c0(r)
where g1 Ficc(r) is the total circulation of the propellerbiade
sec-tions at radius r,
g2r2(r) is that of FGV sections
at
radius r.
Therefore
ut(r) = O
if
g1a F1c(r) + g2 I'2cc(r) = O
(4)
For the case of propulsor with finite number of blades, the
equation
g1F1(r) + g2F2(r) = O
is employed with approximation as guide for defining circulation F2(r),
namely
g1
F2(r) =
(5)
g2
where
g1number of blades of the propeller
g2
number of blades of FGV
F1(r) circulation distribution along propeller blades
F2(r) circulation distribution along
FGVEquation (5) shows that F2(r) depends on F1(r) and ratio of
number
of blades only.
At FGV, the mean tangential interference velocity induced by
fcrward
propeller can be written as
1 r2rr
g1 F1(r)
1t2 =
-urde
=2TTr
(6)
while at forward propeller the mean tangential interference
velocity
induced by FGV is regarded as zero.
Let
pt2 = g
Dz-G1(x)
X 3(7)
ut =
2iTr(3)
ut2
2r
t2
V D2s
then
where F1.(x) Gi(x) -T[D1v D2
X =X
D1D1 and D2 are the diameter of propeller and FGV respectively.
The axial interference velocity is calculated by means of the induc-ed velocity of a sink disc { i} on which the load distribution is
varied along the radius. After making a little modification, the program (Reference to CSSRC Report) for designing contra-rotating propellers can be used to design the combined propulsor in this paper. Fig.2 shows the velocity components at blade section. Then
the hydrodynamic pitch angle will be
uc
Utf
Ua't)arUs I /
tï=Us(1Wt'x))
r'
Fig.2 Velocity polygons
j .J s tj tj 1 w(x)
+ lPaj]+
ua./v -Q.x+
-u/v
sj tj tj s where TÍD.n. iQ.
3j
Sj
X s sj w(x) is wake fractionThence lift-length coefficient and cavitation number of blade element can be calculated. And then the main parameters of the blade
sec-tion are selected based on the strength calculasec-tion and the initial cavitation schemes. In all calculations, the computer programmes derived from those for contra rotating propeller were employed.
Propulsor Models and Test
A pair of combined propulsors was designed with the conditions of:
total thrust required T 106520 kg
y (1.- w(x) + 1 + u aJ aj
tg..(x)
= Dfl.x + -IJQ2 Lis t2t2.J (8)Since the model tests had been planned to he performed in uniform flow in the cavitation tunnel of CSSRC, the propulsor was
design-ed with open water condition. The form of circulation
distribu-tion of the propeller was specified as shown in Fig. 3. Boss ratio is 0.25 for the propeller and 0.261 for FGV. Table i shows the
main parameters of propulsor models from the calculations.
FC X) .0 0.8 0.6
u4
0.2 Q2 04 0.6 0.8 .0Fig.3 Radial distribution to the circula-tion on the propeller blade
In order to compare with Çoflventlonalpropel].er, a single propel-1er without FGV was designed by the method and the computer pro-gram mentioned in 3}. The results are given in the table 2.
Table 1. Geometric Parameters of Blade Section Parameters of the Propeller
rotational speed n 123 RPM
submerged depth of shaft
centre line Hd 8.8 M
advance speed vs 15 knt
diameter of propeller D1 6.17 M
diameter of FGV D2 5.92 m
gap between propeller D 1.3 w
and FGV
number of blades Z 5
s k ew 0°
rake 00
x
(blade area ratio Ad/A
L/D t/D = 0.533) F/L H/D 0.25 0.1456 0.04320 0.0000 0.4888 0.30 0.1676 0.03691 0.01332 0.7496 0.40 0.2073 0.02784 0.02254 0.9784 0.50 0.2412 0.02104 0.02236 1.0300 0.60 0.2691 0.01653 0.01887 1.0100 0.70 0.2911 0.01314 0.01544 0.9726
-5-Table 1. Geometric Parameters of Blade Section Parameters of the Propeller
<blade area ratio Ad/A = 0.533)
x L/D t/D F/L H/D
Parameters of FGV (blade area ratio Ad/A = 0.317)
Table 2 Geometric Parameters of Single Propeller
Sect ion
-6-x L/D t/D F/L H/D 0.261 0.08734 0.02990 0.0000 2.398 0.3001006
0.02585 0.02526 7.895 0.40 0.1244 0.01742 0.05289 25.357 0.50 0.1447 0.01176 0.06275 45.799 0.60 0.1615 0.00890 0.06020 60.500 0.70 0.1747 0.00742 0.05796 64.099 0.80 0.1663 0.00595 0.05621 56.900 0.90 0.1221 0.00447 0.05694 45.158 1.00 0.0000 0.00300 0.0000 29.700 z(blade area ratio Ad/A
L/D t/D 0.502) F/L H/D 0.25 0.1193 0.0491 0.0000 0.833 0.30 0.1476 0.0425 0.02426 0.905 0.40 0.1956 0.0305 0.03058 0.984 0.50 0.2381 0.0227 0.02681 1.008 0.60 0.2647 0.0166 0.02174 0.978 0.70 0.2672 0.0114 0.01835 0.935 0.80 0.2561 0.0086 0.01520 0.867 0.90 0.1870 0.0058 0.01201 0.750 1.00 0.0000 0.0030 0.0000 0.590 0.80 0.2771 0.00976 0.01305 0.09236 0.90 0.2034 0.00638 0.01100 0.08011 1.00 0.0000 0.00300 0.0000 0.04711
A 0.6l A 0.43
p p
Fig. Model test in cavitation tunnel
Fig.4 shows the potographs of the model installation and test in the cavitation tunnel. Because of FGV operating in the slipstream, its
pitch angle is dependent on the induced and interference velocities. Naturally, the difference between theoretical calculation and real
velocity field will lead to deviation in prediction of performance characteristics. Hence during the tests, pitch angle of FGV were ad-justed from the designed position like a controllable pitch propeller. Tests with three pitch angles (O = 00,_40, -8°) and two cavitation numbers were carried out. O = 0° repesents the designed pitch angle position. Negative angle repesents the pitch angle rotated to reduc-ed pitch position.
Results and Analysis
Table 3 gives results of theoretical calculatiQn and test under the design condition.
Table 3 Comparision of Experimental Results and Calculations (A = 0.61) 7 propeller plus FGV single propeller results of CT 1.122 1.1399 Theoretical calculation C 1.778 1.841 fl 0.631 0.619 CT 1.259 1.113 experimental results C 1.953 1.896 flp 0.645 0.592
From above results it seems that results of calculation based on pre-sent computer program are smaller than the results of model tests, while
results of calculation based on the program in 3} are nearer to the test results. As regards comparision of results of two propulsors, the power coefficient of the propeller plus FGV is 3% higher than the single propeller, while thrust coefficient is 13% higher. An improve-ment of 8.9% in efficiency can be achieved compared with the single propeller.
The influence of FGV pitch angle on performance characteristics is given in Table 4. As can be readily seen by this table, the reduction in the pitch angle causes total thrust coefficient to increase, but no improvement in efficiency can be obtained . While under the design condition,
Table 4 Comparision of Propeller with FGV
and conventional Propeller = 0.61)
the thrust coefficient is increased by 2.7%(when e=-4°) ,propeller
effici-ency reaches an improvement of 6.7%, provided that gap between the propel-ler and FGV is reduced by 9% in testing. The influence of gap on per-formance characteristics is described in Fig.5. For the sake of com-paring this results with one of the B-series propellers and AU-series propellers at the same advance coefficient,
0.6
05
0.4 0.3 Q2 0. 00.4-O- °D=62
Dginm
05
0.6 -8-07 B-series propelirr 0.8>\Fig.5 Comparison of Performance characteristics for different pitch angle
Au-s en. es propeller KT 0.184 0.192
0.l7
0.191 0. 177 0. 184 KQ 0.0277 0.0279 0.0291 0.0289 0.0279 0.0279 0.645 0.668 0.626 0.640 0.616 0.641 propeller plus FGV e =O=
4°=-4°
=-8°
.those propellers are designed by means of equivalent torque method as shown in the table. Under the condition of equal horsepower to be absorbed by the propulsors, the propeller with FGV achieves an improvement in efficiency by 8.4% as compared with B-series propel-lers, 4.2% compared with AU-series propellers for O*_40.
Table 5. Thrusts of Propeller arid FGV = 0.61)
Table 5 shows that total thrust of the full scale combined
propul-sor. In the optimal case, the FGV thrust reaches 4. 5° of the total
thrust.
Figure 6 and 7 show that variation of FGT spindle tornue coeffici-ent C and torque coefficient CQ for different pitch angle.It can he seen from these figure that variation in Cq and CQ is
remarka-ble. Under the design condition, the spindle torque coefficient
for _40 is minimum, torque coefficient for P = -S0 is minimum.
Thrust coefficients of FG" for different pitch angle P are given i.n
Figure 8. From this it can he seen that thrust coefficient is
in-creased with reduction in pitch angle. And the thrust tends to in-crease rapidly toward with decreasing in advance coefficient
X.
Reduction in the gap between two propulsors will leac to incrementin the thrust coefficient obviously.
30
20
0.4 0.5 0.6 0.7 0.8
0
0.5 0.6
a7
Q8Xp
0.9Fig.6 Curves of spindle torque coef- Fig.7 Curves of torque coeffici-ficient for different O ent for different P (D'68mm)
(* A* repesents a status in contr?ct'ion of gap by 9%)
0=0
calcul a t i on
0=o°
O=40
O'=-4
e=_0
Total thrust of propulsor T(kg) 104601.3 119016.2 119016.2 122198.4 121562.0 Thrust of FGV
T(kg)
1391.9 1909.4 1877.5 5473.5 3818.7 ratio T /T% w 1.3 1.6 1.6 4.5 3.1CWT
Fig8 Comparision of thrust coefficient for different & and D
This implies that the gap variations have important effect on perform-ance characteristics.
Conclusions
It can he seen from the theoretical and experimental results in the present paper that an improvement in thrust and efficiency will be achieved for combined propulsor. The rotational energy losses in the slipstream can be regained by FGIT installed behind the propeller. If the optimum condition is considered, for the same shaft horsepower, an improvement in efficiency by 8.4% as compared with 8-series and 4.2% as compared with AU-series can he reached.
Fron' the present experimental results, main points are summarized
as
follows:
The adjustment of the pitch angle of FGV (e.g. pitch and its dis-tribution) has some effect on thrust, hut not the efficiency. This means care must be taken to design the FGV for the purpose of in-creasing efficiency. The adjustment of blade ang]e has an important effect on the propulsor torque and spindle torque.
The effect of the gap between propeller and FGV on efficiency
is
obvious. In the present test, thrust coefficient can he increased by 2.7%, efficiency by 6.7%, provided that the gap is decreased by
9% The thrust of FGV may be increased remarkably with decreasing ad-vance coefficient 7. The effect of FGV
IS
Somewhat similar to that of nozzle of the ducted propeller, the thrust produced hy FGV is
de-pendant on the propeller loading. For heavily loaded condition,
the
installation of the FGV will be advantageous, because high energy loss in the slipstream can be recovered by FGV.
Acknowledgments
Authors wish to extend their thanks to Mr. Wu Yiuhwa for his enthusi-astic help in performing the model tests in the cavitation tunnel
of CssRc.
-Symbols ITS P nD d aj F/L T T 71
22
-flfl y 8 s C P lT22
PD
V 8 s np T Cw =pvS
TMQ ÇDvZS.h CMS =Qv Sh
advance coefficient of propeller
distance between penerator line of two proneller
radial distribution of blade
circu-lat ion
blade number
u tangential interference
velocities ti
tj coefficient of tangential
inter-ference velocities
coefficient of axial interference
velocit ies
TtDn-j
reciprocal of adance ratio )\
S3 V s w(x) wake fraction L/D chord-jap'eter ratio t/P ti'ickness-diaaeter ratio HIT) pitch-diameter ratio camber-chord ratio thrust coefficient power coefficient propeller efficiency
thrust coefficient f fixed vanes
torque coefficient of fixed vanes
spindle torque coeff of fixed vanes
p
water density
s
expanded area of blade
b
even width of blade
rotating angle of blade
i
subscript variable of propel] er code
number, li
12
-j
= 1 for the forward propeller j = 2 for the fixed vanesReference
Dong Shi-Tang, 19Th: "Theoretical calculation method of contra-rotating propeller." CSSRC Reprot.
Otto Grim: "Propeller and vane wheel" JSR Vol.24, No.4, Dec. 1980
Huang Rudoa and Cheng Zeiyuan, 19x2: "Propeller design and cal-culation method considered non-uniform inflow field" CSSRC Report.
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