Journal bearings
Chapter Outline
5.1 Introduction 169
5.2 Sliding bearings 171
5.2.1 Lubricants 175
5.3 Design of boundary-lubricated bearings 176
5.4 Design of full film hydrodynamic bearings 179
5.4.1 Reynolds equation derivation 184
5.4.2 Design charts for full-film hydrodynamic bearings 199
5.4.3 Alternative method for the design of full film hydrodynamic bearings 219
5.5 Conclusions 227
References 227
Websites 228
Standards 228
Further reading 230
Nomenclature
Generally, preferred SI units have been stated.
a
radius (m)
b
outer radius (m)
bo
characteristic length (m)
c
radial clearance (m)
D
journal diameter (m)
e
eccentricity
Eu
Euler number (dimensionless)
f
coefficient of friction
Fg
geometrical factor
Fr
Froude number (dimensionless)
Fx
body force component in the x direction (N/m3)
body force component in the y direction (N/m3)
Fy
Fz
body force component in the z direction (N/m3)
g
acceleration due to gravity (m/s2)
h
film thickness (m)
minimum film thickness (m)
ho
Ic
cosine load integral
Is
sine load integral
L
characteristic length, length (m), bearing length (m)
characteristic length (m)
lo
N
speed (normally in rpm)
Mechanical Design Engineering Handbook. https://doi.org/10.1016/B978-0-08-102367-9.00005-6
© 2019 Elsevier Ltd. All rights reserved.
5
168
Ns
p
pmax
P
qx
qy
Q
Qπ
Qs
r
rm
R
Re
Rex
Rey
Rez
S
SAE
t
to
t*
T
Tav
T1
T2
Ta
Tam
Tam,cr
ur
ux
ux,o
uy
uy,o
uz
uz,o
uϕ
u
Uo
V
W
x
x*
y*
z
z*
α
ε
ϕpmax
ϕho
Mechanical Design Engineering Handbook
journal speed (revolutions per second)
static pressure (Pa)
maximum annulus static pressure (N/m2)
load capacity (N/m2), load (N)
volumetric flow rate per unit width in the x direction (m2/s)
volumetric flow rate per unit width in the y direction (m2/s)
lubricant flow rate (m3/s)
flow through the minimum film thickness (m3/s)
side flow rate (m3/s)
radius (m)
mean radius (m)
specific gas constant (J/kg K)
Reynolds number (dimensionless)
Reynolds number based on x component of velocity (dimensionless)
Reynolds number based on y component of velocity (dimensionless)
Reynolds number based on z component of velocity (dimensionless)
geometrical factor, bearing characteristic number
Society of Automotive Engineers
time (s)
characteristic time scale (s)
dimensionless time ratio (dimensionless)
temperature (°C, K)
average temperature (°C)
temperature of the lubricant supply (°C)
temperature of the lubricant leaving the bearing (°C)
Taylor number (dimensionless)
Taylor number based on mean radius (dimensionless)
critical Taylor number based on mean annulus radius (dimensionless)
velocity component in the r direction (m/s)
velocity component in the x direction (m/s)
characteristic velocity component in the x direction (m/s)
velocity component in the y direction (m/s)
characteristic velocity component in the y direction (m/s)
velocity component in the z direction (m/s)
characteristic velocity component in the z direction (m/s)
velocity component in the tangential direction (m/s)
average velocity (m/s)
characteristic velocity (m/s), reference velocity (m/s)
journal velocity (m/s); velocity (m/s)
applied load (N)
distance along the x axis (m),
dimensionless location (dimensionless)
dimensionless location (dimensionless)
distance along the z axis (m)
dimensionless axial location (dimensionless)
angle (degree or rad)
eccentricity ratio
angular position of maximum pressure (degree)
angular position of minimum film thickness (degree)
Journal bearings
ϕpo
ϕ
169
ω
Ω
film termination angle (degree)
position of the minimum film thickness (degree), attitude angle (degree),
azimuth angle (rad)
relative azimuth angle (rad)
nondimensional circumferential location (dimensionless)
viscosity (Pa s)
reference viscosity (Pa s)
dimensionless viscosity (dimensionless)
kinematic viscosity (m2/s)
density (kg/m3)
dimensionless density (dimensionless)
reference density (kg/m3)
squeeze number (dimensionless)
viscous shear stress acting in the x direction, on a plane normal to the z direction
(N/m2)
viscous shear stress acting in the y direction, on a plane normal to the z direction
(N/m2)
angular velocity (rad/s)
angular velocity magnitude (rad/s)
5.1
Introduction
ϕ0
ϕ*
μ
μo
μ*
ν
ρ
ρ*
ρo
σs
τ zx
τ zy
The purpose of a bearing is to support a load, typically applied to a shaft, whilst allowing relative motion between two elements of a machine. The two general classes of
bearings are journal bearings, also known as sliding or plain surface bearings, and
rolling element bearings, sometimes also called ball-bearings. The aims of this chapter
are to describe the range of bearing technology, to outline the identification of which
type of bearing to use for a given application, and to introduce journal bearing design
with specific attention to boundary lubricated bearings and full film hydrodynamic
bearings. The selection and use of rolling element bearings is considered in Chapter 6.
The term ‘bearing’ typically refers to contacting surfaces through which a load is
transmitted. Bearings may roll or slide or do both simultaneously. The range of bearing types available is extensive, although they can be broadly split into two categories:
sliding bearings also known as journal or plain surface bearings, where the motion is
facilitated by a thin layer or film of lubricant, and rolling element bearings, where the
motion is aided by a combination of rolling motion and lubrication. Lubrication is
often required in a bearing to reduce friction between surfaces and to remove heat.
Fig. 5.1 illustrates two of the more commonly known bearings: a journal bearing
and a deep groove ball bearing. A general classification scheme for the distinction
of bearing types is given in Fig. 5.2.
As can be seen from Fig. 5.2 the scope of choice for a bearing is extensive. For a
given application it may be possible to use different bearing types. For example in a
small gas turbine engine rotating as say 50,000 rpm, either rolling bearings or journal
bearings could typically be used although the optimal choice will depend on a number
of factors such as life, cost and size. Fig. 5.3 can be used to give guidance for which
kind of bearing has the maximum load capacity at a given speed and shaft size and
170
Mechanical Design Engineering Handbook
Journal
bearing
Deep groove
ball bearing
Fig. 5.1 A journal bearing and a deep grove ball bearing.
Bearings for rotary motion
Sliding or journal bearings
Rolling bearings
Rolling element
Sliding bearings
Ball
Metallic
deep groove
spherical
self-aligning
Nonmetallic
Roller
Needle
Cylindrical
Taper
(Loading)
Axial
Boundary lubricated
(Lubrication)
(or support)
Radial
(Configuration)
Solid bushing
Mixed film
Barrel
Split bushing
Hydrodynamic
Hydrostatic
Magnetic
Fig. 5.2 Bearing classification.
Based on a taxonomy originally developed by Hindhede, U., Zimmerman, J.R., Hopkins, R.B.,
Erisman, R.J, Hull, W.C., Lang, J.D. Machine Design Fundamentals. A Practical Approach.
Wiley, 1983.
Table 5.1 gives an indication of the performance of the various bearing types for some
criteria other than load capacity.
This and the following chapter provide an introduction to bearings. Lubricant
film sliding bearings are introduced in Section 5.2, the design of boundary lubricated
Journal bearings
171
10,000,000
=5
00
D
mm
x
pro
Ap
D = 250
1,000,000
Hyd
r
bea odynam
ring
ic
s
D = 25
0 mm
m
m
Rolling
elem
bearing ent
s
D = 500
mm
xs
ma
D = 10
0m
m
co
D = 50
erc
50
ial
Approximate
maximum speed
of gas turbine
special ball
bearings
D
D=
10
mm
eel
mm
d st
D = 5 mm
m
soli
0 mm
ate
s
ng
ari
D=1
D=5
=2
5m
be
D = 10 mm
mm
xim
g
lin
rol
5 mm
ro
App
D=2
10,000
D=
mm
mm
D = 25 mm
1000
10
of
Typical maximum load (N)
D = 50 mm
ed
100,000
D=
pe
0 mm
=5
m
m
mit
rst li
Rubbing
plain
bearings
ft bu
sha
D
100
Plain bearings
10
0.01
0.1
1
10
100
1000
10,000
100,000
Speed (Revolutions per second)
Fig. 5.3 Bearing type selection by load capacity and speed.
Reproduced with permission from Neale, M.J. (ed.), 1995. The Tribology Handbook.
Butterworth Heinemann.
bearings which are typically used for low speed applications are considered in
Section 5.3, the design of full-film hydrodynamically lubricated bearings is described
in Section 5.4 and Chapter 6 considers rolling element bearings. For further reading
the texts by Khonsari and Booser (2017), Harris (2001) and Brandlein et al. (1999)
provide an extensive overview of bearing technologies.
5.2
Sliding bearings
The term sliding bearing refers to bearings where two surfaces move relative to each
other without the benefit of rolling contact. The two surfaces slide over each other and
this motion can be facilitated by means of a lubricant which gets squeezed by the
motion of the components and can generate sufficient pressure to separate them,
thereby reducing frictional contact and wear.
A typical application of sliding bearings is to allow rotation of a load-carrying
shaft. The portion of the shaft at the bearing is referred to as the journal and the stationary part, which supports the load, is called the bearing (Fig. 5.4). For this reason,
172
Table 5.1 Comparison of bearing performance for continuous rotation
Combined axial and
radial load
capability
Low starting
torque
capability
Silent
running
Standard
parts
available
Lubrication
simplicity
Rubbing plain
bearings
(nonmetallic)
Porous metal plain
bearings oil
impregnated
Fluid film
hydrodynamic
bearings
Hydrostatic bearings
Poor
Some in most cases
Poor
Fair
Some
Excellent
Good
Some
Good
Excellent
Yes
Excellent
Fair
No; separate thrust
bearing needed
Good
Excellent
Some
Excellent
Excellent
Excellent
No
Rolling bearings
Good
No; separate thrust
bearing needed
Yes in most cases
Very good
Usually
satisfactory
Yes
Usually requires a
recirculation
system
Poor special
system needed
Good when grease
lubricated
Bearing type
Reproduced with permission from Neale, M.J. (ed.), 1995. The Tribology Handbook. Butterworth Heinemann.
Mechanical Design Engineering Handbook
Accurate
radial
location
Journal bearings
173
Shaft
Journal
diameter
Bearing
diameter
Journal
Journal
Clearance
Bearing
material
Housing block
Fig. 5.4 A plain surface, sliding or journal bearing.
sliding bearings are often collectively referred to as journal bearings, although this
term ignores the existence of sliding bearings that support linear translation of components. Another common term frequently used in practice is plain surface bearings.
This section is principally concerned with bearings for rotary motion and the terms
journal and sliding bearing are used interchangeably.
There are three regimes of lubrication for sliding bearings:
(i) boundary lubrication,
(ii) mixed film lubrication,
(iii) full film lubrication.
Boundary lubrication typically occurs at low relative velocities between the journal
and the bearing surfaces and is characterised by actual physical contact. The surfaces,
even if ground to a low value of surface roughness, will still consist of a series of peaks
and troughs as illustrated schematically in Fig. 5.5. Although some lubricant may
be present, the pressures generated within it will not be significant and the relative
motion of the surfaces will bring the corresponding peaks periodically into contact.
Mixed film lubrication occurs when the relative motion between the surfaces is
Journal surface
Boundary
lubrication
Bearing surface
Journal surface
Mixed-film
lubrication
Bearing surface
Journal surface
Full film hydrodynamic
lubrication
Bearing surface
Fig. 5.5 A schematic representation of the surface roughness for sliding bearings and the
relative position depending on the type of lubrication occurring.
174
Mechanical Design Engineering Handbook
sufficient to generate high enough pressures in the lubricant film, which can partially
separate the surfaces for periods of time. There is still contact in places around the
circumference between the two components. Full film lubrication occurs at higher
relative velocities. Here the motion of the surface generates high pressures in the lubricant, which separate the two components and journal can ‘ride’ on a wedge of fluid.
All of these types of lubrication can be encountered in a bearing without external
pressuring of the bearing. If lubricant under high enough pressure is supplied to
the bearing to separate the two surfaces it is called a hydrostatic bearing (Rowe, 2012).
The performance of a sliding bearing differs markedly depending on which type of
lubrication is physically occurring. This is illustrated in Fig. 5.6, which shows the
variation of the coefficient of friction with a group of variables called the bearing
parameter which is defined by:
μN
P
(5.1)
where:
μ ¼ viscosity of lubricant (Pa s),
N ¼ speed (for this definition normally in rpm),
P ¼ load capacity (N/m2) given by
P¼
W
LD
(5.2)
W ¼ applied load (N),
L ¼ bearing length (m),
D ¼ journal diameter (m).
The bearing parameter, μN/P, groups several of the bearing design variables into one
number. Normally, of course, a low coefficient of friction is desirable to ensure that
only a small amount of power is necessary to rotate or drive the component concerned.
Fig. 5.6 Schematic representation of
the variation of bearing performance
with lubrication.
Coefficient of friction, F
II
I
Boundary
lubrication
II
Mixed film
lubrication
III
Hydrodynamic
lubrication
I
III
Bearing
parameter
µN/P
Journal bearings
175
Typical coefficients of friction for a boundary lubricated bearing are between approximately 0.05 and 0.1. By contrast the coefficient of friction for a rolling element bearing is typically of the order of 0.005.
In general, boundary lubrication is used for slow speed applications where the surface speed is less than approximately 1.5 m/s. Mixed-film lubrication is rarely used
because it is difficult to quantify the actual value of the coefficient of friction and
small changes in, for instance, the viscosity will result in large changes in friction
(note the steep gradient in Fig. 5.6 for this zone). The design of boundary-lubricated
bearings is outlined in Section 5.3 and full-film hydrodynamic bearings in Section 5.4.
5.2.1
Lubricants
As can be seen from Fig. 5.6 bearing performance is dependent on the type of lubrication occurring and the viscosity of the lubricant. The viscosity is a measure of a
fluid’s resistance to shear. Lubricants can be solid, liquid or gaseous although the most
commonly known are oils and greases. The principal classes of liquid lubricants are
mineral oils and synthetic oils. Their viscosity is highly dependent on temperature as
illustrated in Fig. 5.7. They are typically solid at –35°C, thin as paraffin at 100°C and
burn above 240°C. Many additives are used to affect their performance. For example
EP (extreme pressure) additives add fatty acids and other compounds to the oil, which
attack the metal surfaces to form ‘contaminant’ layers, which protect the surfaces and
10,000
2000
1000
300
200
Viscosity (mPa s)
100
40
30
SA
E
SA 60
E
5
SA 0
E
4
SA 0
E
30
SA
E
20
20
10
SAE 60, ISO VG 320
SAE 50, ISO VG 220
SAE 40, ISO VG 150
SAE 30, ISO VG 100
SA
7
6
5
E
IS
O
IS
O
4
10
VG
VG
32
SAE 20, ISO VG 68
22
3
SAE 10, ISO VG 46
ISO VG 32
ISO VG 22
2
10
20
30
40
50
60
70
80 90 100
130
Temperature (°C)
Fig. 5.7 Variation of absolute viscosity with temperature for various lubricants.
176
Mechanical Design Engineering Handbook
reduce friction even when the oil film is squeezed out by high contact loads. Greases
are oils mixed with soaps to form a thicker lubricant that can be retained on surfaces.
The viscosity variation with temperature of oils has been standardised and oils are
available with a class number, for example SAE 10, SAE 20, SAE 30, SAE 40, SAE
5 W and SAE 10 W. The origin of this identification system developed by the Society
of Automotive Engineers was to class oils for general-purpose use and winter use, the
‘W’ signifying the latter. The lower the numerical value, the thinner or less viscous the
oil. Multigrade oil (e.g. SAE 10 W/40) is formulated to meet the viscosity requirements of two oils giving some of the benefits of the constituent parts. An equivalent
identification system is also available from the International Organisation for Standardisation (ISO 3448).
5.3
Design of boundary-lubricated bearings
As described in Section 5.2, the journal and bearing surfaces in a boundary lubricated
bearing are in direct contact in places. These bearings are typically used for very low
speed applications such as bushes and linkages where their simplicity and compact
nature are advantageous. Examples include household lawnmower wheels, garden
hand tools such as shears, ratchet wrenches and domestic, automotive door hinges
incorporating the journal as a solid pin riding inside a cylindrical outer member
and even for cardboard wheels (Fig. 5.8). A further example, involving linear motion
is the slider in an ink jet printer.
General considerations in the design of a boundary lubricated bearing are:
l
l
l
l
l
l
the coefficient of friction (both static and dynamic),
the load capacity,
the relative velocity between the stationary and moving components,
the operating temperature,
wear limitations and
the production capability.
Fig. 5.8 The Move-it wheel for box transportation—Courtesy of Move-it.
Journal bearings
177
A useful measure in the design of boundary-lubricated bearings is the PV factor (load
capacity peripheral speed), which indicates the ability of the bearing material to
accommodate the frictional energy generated in the bearing. At the limiting PV value
the temperature will be unstable and failure will occur rapidly. A practical value for
PV is half the limiting PV value. Values for PV for various bearing materials are given
in Table 5.2, and Lancaster (1971), Yamaguchi and Kashiwagi (1982), and Koring
(2012). The preliminary design of a boundary-lubricated bearing essentially consists
of setting the bearing proportions, its length and its diameter, and selecting the bearing
material such that an acceptable PV value is obtained. This approach is set out step-bystep in Table 5.2.
Example 5.1. A bearing is to be designed to carry a radial load of 700 N for a shaft of
diameter 25 mm running at a speed of 75 rpm (Fig. 5.9). Calculate the PV value and by
comparison with the available materials listed in Table 5.3 determine a suitable bearing material.
Solution
The primary data are W ¼ 700 N, D ¼ 25 mm and N ¼ 75 rpm.
Use L/D ¼ 1 as an initial suggestion for the length to diameter ratio for the bearing.
L ¼ 25 mm.
Calculating the load capacity, P
W
700
¼
¼ 1:12 MN=m2
LD 0:025 0:025
2π
D
0:025
N ¼ 0:1047 75
¼ 0:09816 m=s
V ¼ ωr ¼
60
2
2
P¼
PV ¼ 0:11 MN=m2 ðm=sÞ
Multiplying this by a factor of 2 gives PV ¼ 0.22 (MN/m s).
A material with PV value greater than this, such as filled PTFE (limiting PV value
up to 0.35), or PTFE with filler bonded to a steel backing (limiting PV value up to
1.75), would give acceptable performance.
Example 5.2. The paper feed for a photocopier is controlled by two rollers which are
sprung together with a force of approximately 20 N. The rollers each consist of a
20 mm outer diameter plastic cylinder pressed onto a 10 mm diameter steel shaft
(Fig. 5.10). The maximum feed rate for the copier is 30 pages per minute. Select bearings to support the rollers (After IMechE, 1994).
Solution
The light load and low speed, together with the requirement that the bearings should
be cheap, clean and maintenance free suggest that dry rubbing bearings be considered.
The length of A4 paper is 297 mm.
The velocity of the roller surface is given by:
Vroller ¼
30 0:297
¼ 0:1485ms1 :
60
178
Mechanical Design Engineering Handbook
Table 5.2 Step-by-step guideline for boundary lubricated bearing design
Start Selection
Determine rotational speed
Determine the speed of rotation of the
bearing. This information may set by the
designer or defined by the specification.
OK
Determine the load to be supported
OK
No
Select L/D
OK
Determine the load to be supported. This
information may set by the designer or
defined by the specification.
No
Calculate the load capacity, P
Set the bearing proportions. Common
practice is to set the length to diameter
ratio between 0.5 and 1.5. If the diameter
is known as an initial trial set L equal to D
Calculate the load capacity, P=W/LD.
OK
Calculate V
Determine the maximum tangential speed
of the journal, V=w r.
OK
Calculate the PV factor
Calculate PV
OK
Select a factor of safety, SF
OK
Determine a factor of safety (SF) based on
knowledge of certainty of the operating
conditions. For general uncertainty a
factor of safety of 2 is normally adequate.
Multiply the PV factor obtained by the
factor of safety
Calculate SF PV
OK
Select a bearing material
OK
Complete
No
Interrogate manufacturer’s data for
bearing materials to identify an
appropriate bearing material with a value
of PV greater than (SF PV)
Journal bearings
179
Ø25
700 N
Fig. 5.9 Boundary lubricated bearing design example (dimensions in mm).
The angular velocity of the rollers is 0.1485/0.01 ¼ 14.85 rad/s.
The velocity of the bearing surface is therefore.
V ¼ 14:85 0:005 ¼ 0:07425m=s:
An L/D ratio of 0.5 could be suitable giving a bearing length of 5 mm. The load capacity is:
W
20
¼
¼ 0:4 MPa::
LD 0:005 0:01
PV ¼ 0:4 0:07425 ¼ 0:0297 MNm2 ms1 :
P¼
For this low value of PV the cheapest bearing materials, thermoplastics are
adequate.
5.4
Design of full film hydrodynamic bearings
In a full film hydrodynamic bearing, the load on the bearing is supported on a continuous film of lubricant so that no contact between the bearing and the rotating journal
occurs. The motion of the journal inside the bearing creates the necessary pressure to
support the load, as indicated in Fig. 5.11. Hydrodynamic bearings are commonly
found in internal combustion engines for supporting the crankshaft, and in turbocharger applications to support the rotor assembly (Fig. 5.12). Hydrodynamic bearings
can consist of a full circumferential surface or a partial surface around the journal
(Fig. 5.13).
Bearing design can involve so many parameters that it is a challenge to develop an
optimal solution. One route towards this goal would be to assign attribute points to
each aspect of the design and undertake an optimisation exercise using multiobjective
design optimisation or a similar scheme (Hashimoto and Matsumoto, 2001; Hirania
and Suh, 2005). This can however be time consuming if the software was not already
in a developed state and would not necessarily produce an optimum result due to inadequacies in modelling and incorrect assignment of attribute weightings. An alternative
180
Table 5.3 Characteristics of some rubbing materials
Material
Limiting
PV value
(MN/m S)
Maximum
operating
temperature
(°C)
1.4–2
0.11
350–500
3.4
0.145
130–350
70
0.28–0.35
2
Coefficient
of friction
Coefficient
of expansion
(×1026/°C)
Comments
Typical
application
For continuous
dry operation
Food and textile
machinery
0.1–0.25
dry
0.1–0.35
dry
2.5–5.0
350–600
0.1–0.15
dry
12–13
0.35
250
0.13–0.5
dry
3.5–5
35
0.35
200
0.1–0.4 dry
25–80
Roll-neck
bearings
10
0.035
100
0.1–0.45
dry
100
Bushes and
thrust washers
10–14
0.035–0.11
100
0.15–0.4
dry
80–100
Bushes and
thrust washers
4.2–5
Suitable for sea
water operation
Mechanical Design Engineering Handbook
Carbon/
graphite
Carbon/
graphite with
metal
Graphite
impregnated
metal
Graphite/
thermosetting resin
Reinforced
thermosetting
plastics
Thermoplastic
material
without filler
Thermoplastic with
filler or
metal backed
Maximum
load
capacity
P(MN/m2)
140
0.35
105
0.2–0.35
dry
27
With initial
lubrication
only
Filled PTFE
7
Up to 0.35
250
0.05–0.35
dry
60–80
Glass, mica,
bronze,
graphite
PTFE with
filler,
bonded to
steel backing
140
Up to 1.75
280
0.05–0.3
dry
20
Woven
PTFE
reinforced
and bonded
to metal
backing
420
Up to 1.6
250
0.03–0.3
–
Sintered bronze
bonded to steel
backing
impregnated
with PTFE/lead
Reinforcement
may be
interwoven
glass fibre or
rayon
For conditions of
intermittent
operation or
boundary
lubrication (e.g.
ball joints,
suspension,
steering)
For dry
operations where
low friction and
wear required
Aircraft controls,
linkages,
gearbox, clutch,
conveyors,
bridges
Aircraft and
engine controls,
linkages, engine
mountings,
bridge bearings
Journal bearings
Thermoplastic
material with
filler bonded
to metal back
Reproduced with permission from Neale, M.J. (ed.), 1995. The Tribology Handbook. Butterworth Heinemann.
181
182
Mechanical Design Engineering Handbook
Steel shaft
Plastic cylinder
Fig. 5.10 Copier rollers.
Reproduced from Institution of Mechanical Engineers, 1994. Tribological Design Data. Part 1:
Bearings. I. Mech. E.
W
N
e
r
h
Øp
o
Øho
o
Øpmax
Film
pressure, p
pmax
Fig. 5.11 Motion of the journal generates pressure in the lubricant separating the two surfaces.
Beyond ho, the minimum film thickness, the pressure terms go negative and the film is ruptured.
Fig. 5.12 Turbocharger journal.
Journal bearings
183
Fig. 5.13 Partial surface
journal bearings.
approach which is sensible as a starting point and outlined here is to develop one or a
number of feasible designs and use judgement to select the best, or combine the best
features of the proposed designs.
The design procedure for a journal bearing, recommended here as a starting point,
includes specification of the:
l
l
l
l
l
l
journal radius, r;
radial clearance, c;
axial length of the bearing surface, L;
type of lubricant and its viscosity, μ;
journal speed, N;
load, W.
Values for the speed, the load and possibly the journal radius are usually specified by
the machine requirements, and stress and deflection considerations. As such, journal
bearing design consists of the determination of the radial clearance, the bearing length
and the lubricant viscosity. The design process for a journal bearing is usually iterative. Trial values for the clearance, the length and the viscosity are chosen, various
performance criteria calculated and the process repeated until a satisfactory or
optimised design is achieved. Criteria for optimisation may be minimising of the
frictional loss, minimising the lubricant temperature rise, minimising the lubricant
supply, maximising the load capability, and minimising production costs.
The clearance between the journal and the bearing depends on the nominal diameter of the journal, the precision of the machine, surface roughness and thermal
expansion considerations. An overall guideline is for the radial clearance, c, to be
in the range 0.001r < c < 0.002r, where r is the nominal bearing radius
(0.001D < 2c < 0.002D). Fig. 5.14 shows values for the recommended diametral
clearance (2 c) as a function of the journal diameter and rotational speed for steadily
loaded bearings.
For a given combination of a, c, L, μ, N and W, the performance of a journal bearing
can be calculated. This requires determining the pressure distribution in the bearing,
the minimum film thickness ho, the location of the minimum film thickness θpmax, the
coefficient of friction f, the lubricant flow Q, the maximum film pressure pmax and
the temperature rise ΔT of the lubricant.
The pressure distribution in a journal bearing (Fig. 5.11) can be determined by solving the relevant form of the Navier–Stokes fluid flow equations, which in the reduced
form for journal bearings is called the Reynolds equation, Eq. (5.3).
∂ h3 ∂p
∂ h3 ∂p
∂h
∂V
+
¼ 6V + 6h
∂x μ ∂x
∂z μ ∂z
∂x
∂x
(5.3)
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Mechanical Design Engineering Handbook
22
6000
5000
250 µm
5
20
0
3000
µm
4000
µm
17
5
µm
15
2000
0
µm
µ
75
µm
0
10
µm
800
700
600
500
50
m
Speed (rpm)
µm
5
12
1000
400
300
200
25 µm
100
80
70
60
25
50
75
100
125
150
175
200
Journal: diameter (mm)
Fig. 5.14 Minimum recommended values for the diametral clearance (2 c) for steadily
loaded journal bearings.
Reproduced with permission from Welsh, R.J., 1983. Plain Bearing Design Handbook.
Butterworth.
where:
h ¼ film thickness (m);
μ ¼ dynamic viscosity (Pa s);
p ¼ fluid pressure (Pa);
V ¼ journal velocity (m/s);
x, y, z ¼ rectangular coordinates (m).
The Reynolds equation was first derived by Reynolds (1886) and further developed by
Harrison (1913) to include the effects of compressibility. Here the derivation of the
general Reynolds equation is given in Section 5.4.1 and is developed following the
general outline given by Hamrock (1994) and Childs (2011). Many readers may prefer
to omit detailed consideration of the derivation as there is no need to go through
this for practical design applications and instead proceed directly to Section 5.4.2.
Nevertheless the derivation is included here for reference.
5.4.1
Reynolds equation derivation
The Navier–Stokes equations in a Cartesian coordinate system, for compressible flow
assuming constant viscosity are given by Eqs (5.4)–(5.6). Here x is taken as the coordinate in the direction of sliding, y is the coordinate in the direction of side leakage and
z is the coordinate across the lubricant film.
Journal bearings
185
ρ
2
∂ux
∂ux
∂ux
∂ux
∂p
∂ ux ∂2 ux ∂2 ux
+ ux
+ uy
+ uz
¼ +μ
+
+
+
∂t
∂x
∂y
∂z
∂x2
∂y2
∂z2
∂x
μ ∂ ∂ux ∂uy ∂uz
+
+
+ Fx
∂y
∂z
3 ∂x ∂x
(5.4)
2
∂uy
∂uy
∂uy
∂uy
∂ uy ∂2 uy ∂2 uy
∂p
+
+ ux
+ uy
+ uz
¼ +μ
+
+
∂t
∂x
∂y
∂z
∂x2
∂y2
∂z2
∂y
μ ∂ ∂ux ∂uy ∂uz
+
+
+ Fy
∂y
∂z
3 ∂y ∂x
(5.5)
ρ
ρ
2
∂uz
∂uz
∂uz
∂uz
∂p
∂ uz ∂2 uz ∂2 uz
+ ux
+ uy
+ uz
¼ +μ
+
+
+
∂t
∂x
∂y
∂z
∂x2
∂y2
∂z2
∂z
μ ∂ ∂ux ∂uy ∂uz
+
+
+ Fz
∂y
∂z
3 ∂z ∂x
(5.6)
If the density is assumed constant then the dilation (Eq. 5.7) is zero.
∂ux ∂uy ∂uz
+
+
¼0
∂x
∂y
∂z
(5.7)
The Navier–Stokes equations for constant density and constant viscosity reduce to
2
∂ux
∂ux
∂ux
∂ux
∂p
∂ ux ∂2 ux ∂2 ux
+ ux
+ uy
+ uz
¼ +μ
+ Fx
+
+
∂t
∂x
∂y
∂z
∂x2
∂y2
∂z2
∂x
(5.8)
2
∂uy
∂uy
∂uy
∂uy
∂ uy ∂2 uy ∂2 uy
∂p
+ ux
+ uy
+ uz
¼ +μ
+ 2 + 2 + Fy
ρ
∂t
∂x
∂y
∂z
∂x2
∂y
∂z
∂y
(5.9)
2
∂uz
∂uz
∂uz
∂uz
∂p
∂ uz ∂ 2 uz ∂ 2 uz
+ ux
+ uy
+ uz
¼ +μ
+ 2 + 2 + Fz
ρ
∂t
∂x
∂y
∂z
∂x2
∂y
∂z
∂z
(5.10)
ρ
For conditions known as slow viscous motion, where pressure and viscous terms predominate simplifications are possible for the Navier–Stokes equations making their
solution more amenable to analytical and numerical techniques.
The Navier–Stokes equations can be nondimensionalised to enable a generalised
solution (Childs, 2011). Here the process is repeated to give dimensionless groups
of specific relevance to journal bearings using the characteristic parameters given
in Eqs (5.11)–(5.20).
x∗ ¼
x
lo
where lo is a characteristic length in the x direction.
(5.11)
186
Mechanical Design Engineering Handbook
y
bo
y∗ ¼
(5.12)
where bo is a characteristic length in the y direction.
z
ho
z∗ ¼
(5.13)
where ho is a characteristic length in the z direction.
t∗ ¼
t
to
(5.14)
where to is a characteristic time.
ux* ¼
ux
ux, o
(5.15)
where ux,o is a characteristic velocity in the x direction.
uy* ¼
uy
uy, o
(5.16)
where uy,o is a characteristic velocity in the y direction.
uz* ¼
uz
uz, o
(5.17)
where uz,o is a characteristic velocity in the z direction.
ρ∗ ¼
ρ
ρo
(5.18)
where ρo is a characteristic density.
μ∗ ¼
μ
μo
(5.19)
where μo is a characteristic viscosity.
p∗ ¼
h2o p
μo ux, o lo
(5.20)
Substitution of the above dimensionless parameters, Eqs (5.11)–(5.20) in Eq. (5.4),
gives
Journal bearings
187
2
lo uy,o
lo ∂ux*
∂u
∂u
lo uz, o
∂u
lo g
μo
lo
1 ∂p∗
+ ux* x* +
uy* x* +
uz* x* ¼ 2 ux, o to ∂t∗
∂x∗ bo ux, o
∂y* ho ux,o
∂z∗ ux,o ρo ux,o lo ho ρ∗ ∂x∗
2 μo 1 ∂
∂ux* uy,o lo ∂uy* uz,o lo ∂uz*
2μo 1 ∂
∂ux*
+
+
μ∗
+
+
μ∗
∂x∗ ux, o bo ∂y∗ ux,o ho ∂z∗
ρo ux,o lo ρ∗ ∂x∗
∂x∗
3 ρo ux,o lo ρ∗ ∂x∗
2
2
μo
lo
1 ∂
∂ux* uy, o bo ∂uy*
μo
lo
1 ∂
∂ux* uz,o ho ∂uz*
μ∗
+
μ∗
+
+
ρo ux, o lo bo ρ∗ ∂y∗
∂y∗ ux, o lo ∂x∗
ρo ux,o lo ho ρ∗ ∂z∗
∂z∗ ux, o lo ∂x∗
(5.21)
The relative importance of inertia and viscous forces can be determined by examination of the value of the Reynolds number
ρ ux, o lo
Re ¼ o
μo
(5.22)
The inverse of the Reynolds number occurs throughout Eq. (5.21).
In fluid film lubrication, because of the dominance of the viscous term ∂2ux∗/∂2z∗, a
modified form of the Reynolds number is used, defined for the x component of
velocity by
ρ ux, o h2o
Rex ¼ o
μo lo
(5.23)
and similarly for the y and z directions
ρ uy, o h2o
Rey ¼ o
μ o bo
(5.24)
ρ uz, o ho
Rez ¼ o
μo
(5.25)
The squeeze number is defined by
ρ h2
σs ¼ o o
μo to
(5.26)
Typically in hydrodynamically lubricated journal bearings, viscous forces are much
greater than inertia forces and a typical Reynolds number, using Eq. (5.23), might be
of the order of 1 104.
Substitution of the Reynolds and squeeze numbers into Eq. (5.21), gives
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Mechanical Design Engineering Handbook
∂ux*
∂ux*
∂ux*
∂ux* lo g
1 ∂p∗ 1 ∂
∂ux*
+ Rex ux*
+ Rey uy*
+ Rez uz*
¼ 2 Rex +
μ∗
∂t∗
∂x∗
∂y∗
∂z∗ ux, o
∂z∗
ρ∗ ∂x∗ ρ∗ ∂z∗
2 ho 2 1 ∂
∂ux* uy,o lo ∂uy* uz, o lo ∂uz*
μ∗
+
+
+
∂x∗ ux, o bo ∂y∗ ux, o ho ∂z∗
3 lo ρ∗ ∂x∗
2
2
ho 1 ∂
∂ux* uy, o bo ∂uy*
ho 1 ∂
∂u
1 ∂
uz, o ho ∂uz*
μ∗
+
μ∗ x* +
μ∗
+2
ρ∗ ∂z∗
bo ρ∗ ∂y∗
∂y∗ ux, o lo ∂x∗
lo ρ∗ ∂x∗
∂x∗
ux, o lo ∂x∗
σs
(5.27)
Examination of the order of magnitude of the terms in Eq. (5.27) provides an indication of the relative significance of the terms and which need to considered. The inertia
terms, gravity term and uz,o/ux,o are of order ho/lo. The pressure gradient term and the
first viscous term are of order unity. The remaining viscous terms are of order (ho/lo)2
or (ho/bo)2 and therefore very small in comparison to the other terms. Neglecting terms
of the order of (ho/lo)2 and (ho/bo)2 gives
∂ux*
∂ux*
∂ux*
∂ux*
+ Rex ux*
+ Rey uy*
+ Rez uz*
∂t∗
∂x∗
∂z∗
∂y∗
lo g
1 ∂p∗ 1 ∂
∂ux*
¼ 2 Rex +
μ
ux, o
ρ∗ ∂x∗ ρ∗ ∂z∗ ∗ ∂z∗
σs
(5.28)
Similarly for the y and z components
∂uy*
∂uy*
∂uy*
∂uy*
+ Rex ux*
+ Rey uy*
+ Rez uz*
∂t∗
∂x∗
∂y
∂z∗
∗
∂uy*
bo g
1 ∂p∗ 1 ∂
¼ 2 Rey +
μ
uy, o
ρ∗ ∂y∗ ρ∗ ∂z∗ ∗ ∂z∗
σs
∂p
¼0
∂z∗
(5.29)
(5.30)
Examination of Eqs (5.28)–(5.30) shows that the pressure is a function of x*, y* and t*,
p ¼ f x∗ , y∗ , t∗
(5.31)
The continuity equation can be expressed as
σs
∂p∗
∂ ∂ ∂ + Rex
ρ∗ ux∗ + Rey
ρ∗ uy* + Rez
ρ uz* ¼ 0
∂t∗
∂x∗
∂y∗
∂z∗ ∗
(5.32)
Taylor vortices can be formed in a fluid filled annulus with inner cylinder rotation as
illustrated in Figs 5.15 and 5.16. Vortices involve significant acceleration and deceleration of the flow and therefore dissipation of energy and are normally undesirable in
Journal bearings
189
W
a
b
Fig. 5.15 Taylor vortices.
Taylor vortex
Outer cylinder
Inner cylinder
Fig. 5.16 Taylor vortex cells.
any bearing lubrication Scheme. A check can readily be made to identify if Taylor
vortices are likely to occur. If the Taylor number, Tam, based on the mean annulus
radius (Eq. 5.33) is greater than approximately 41.2, then Taylor vortices may
form and laminar flow conditions no longer hold invalidating the use of the equations
developed in this section.
Tam ¼
Ωrm0:5 ðb aÞ1:5
ν
(5.33)
rm ¼ 0:5ðb + aÞ
(5.34)
pffiffiffiffiffiffiffiffiffiffi
Tam,cr ¼ 1697 ¼ 41:19
(5.35)
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Mechanical Design Engineering Handbook
The critical angular velocity for this case is given by
Ωcr ¼
41:19ν
rm0:5 ðb aÞ1:5
(5.36)
For an annulus with a finite gap,
Ωcr ¼
41:19νFg
rm0:5 ðb aÞ1:5
(5.37)
where Fg is a geometrical factor defined by
π2
b a 1
pffiffiffi 1 Fg ¼
2rm
41:19 S
(5.38)
and S is given in an alternative form by
S ¼ 0:0571 1 0:652
1
ðb aÞ=rm
ðb aÞ=rm
+ 0:00056 1 0:652
1 ðb aÞ=2rm
1 ðb aÞ=2rm
(5.39)
The Froude number (Eq. 5.40) provides an indication of the relative importance of
inertia and gravity forces. A typical Froude number for a journal bearing might be
of the order of 10, providing an indication that gravity forces can be neglected in comparison with viscous forces.
Uo
(5.40)
Fr ¼ pffiffiffiffiffiffi
gL
The importance of pressure relative to inertia can be judged by examination of the
Euler number (Eq. 5.41). For a typical journal bearing the Euler number may be of
the order of 100, giving an indication that the pressure term is much larger than
the inertia term.
Eu ¼
p po
0:5ρUo2
(5.41)
If in addition to neglecting terms of order (ho/lo)2 or (ho/bo)2, terms of the order of ho/lo
and ho/bo are neglected and only terms of the order of unity are considered, the
Navier–Stokes equations reduce to
∂p ∂
∂ux
¼
μ
∂z
∂x ∂z
(5.42)
∂uy
∂p ∂
¼
μ
∂z
∂y ∂z
(5.43)
Journal bearings
191
Eq. (5.31) shows that the pressure is only a function of x and y for steady state conditions. Eqs (5.42), (5.43) can therefore be integrated directly to give general expressions for the velocity gradients as follows.
∂ux z ∂p A
¼
+
∂z μ ∂x μ
(5.44)
∂uy z ∂p C
¼
+
∂z μ ∂y μ
(5.45)
where A and C are constants of integration.
The temperature across the thin layer of layer lubricant in a journal bearing may
vary significantly. As viscosity is highly dependent on temperature, this leads to
increased complexity in obtaining a solution to Eqs 5.44 and 5.45. It has been found
acceptable, however, in many fluid film applications to model the viscosity of a fluid
film using an average value for the viscosity across the film. With μ taken as the average value of viscosity across the fluid film, then integration of Eqs 5.44 and 5.45 gives
the velocity components as follows.
ux ¼
z2 ∂p
z
+A +B
2μ ∂x
μ
(5.46)
uy ¼
z2 ∂p
z
+C +D
2μ ∂y
μ
(5.47)
where B and D are constants of integration.
If the no-slip condition is assumed at the fluid solid interface then the boundary
conditions are as follows and illustrated in Fig. 5.17.
ux ¼ ux, a , uy ¼ uy, a at z ¼ 0
ux ¼ ux, b , uy ¼ uy, b at z ¼ h
Application of the boundary conditions to the equations for the velocity gradients and
velocity components (Eqs 5.46 and 5.47) gives
∂ux 2z h ∂p ux, a ux, b
¼
∂z
h
2μ ∂x
(5.48)
∂uy 2z h ∂p uy, a uy, b
¼
∂z
h
2μ ∂y
(5.49)
ux ¼
zðz hÞ ∂p h z
z
+
ux, a + ux, b
2μ ∂x
h
h
(5.50)
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Mechanical Design Engineering Handbook
uy, b
uy, a
b
ux, b
ux, a
a
a
b
Fig. 5.17 Journal bearing boundary conditions.
uy ¼
zðz hÞ ∂p h z
z
+
uy, a + uy, b
2μ ∂y
h
h
(5.51)
The viscous shear stresses are defined by
∂uz ∂ux
τzx ¼ μ
+
∂x
∂z
(5.52)
∂uz ∂uy
+
τzy ¼ μ
∂y
∂z
(5.53)
The order of magnitude of ∂ uz/∂ x and ∂ uz/∂ y are much smaller than ∂ ux/∂ z and ∂ uy/∂ z
so the viscous shear stresses can be approximated by
τzx ¼ μ
∂ux
∂z
(5.54)
τzy ¼ μ
∂uy
∂z
(5.55)
From Eqs (5.48), (5.49), the viscous shear stresses acting on the solid surfaces can be
expressed by
∂ux
h ∂p μðux, a ux, b Þ
ðτzx Þz¼0 ¼ μ
¼
∂z z¼0
2 ∂x
h
(5.56)
Journal bearings
193
∂ux
h ∂p μðux, a ux, b Þ
+
ðτzx Þz¼h ¼ μ
¼
∂z z¼h
2 ∂x
h
(5.57)
∂uy
h ∂p μ uy, a uy, b
¼
τzy z¼0 ¼ μ
∂z z¼0
2 ∂y
h
(5.58)
∂uy
h ∂p μ uy, a uy, b
¼
+
τzy z¼h ¼ μ
∂z z¼h
2 ∂y
h
(5.59)
The negative signs for the viscous shear stresses in Eqs (5.56)–(5.59) indicate that the
stress acts in a direction opposite to the motion.
The volumetric flow rates per unit width in the x and y directions are defined by
qx ¼
Z h
ux dz
(5.60)
uy dz
(5.61)
0
qy ¼
Z h
0
Substituting for the velocity components using Eqs (5.50), (5.51) in Eqs (5.60), (5.61)
gives
qx ¼ h3 ∂p ux, a + ux, b
h
+
12μ ∂x
2
(5.62)
qy ¼ h3 ∂p uy, a + uy, b
h
+
12μ ∂y
2
(5.63)
The Reynolds equation is formed by substituting the expressions for the volumetric
flow rate into the continuity equation.
Integrating the continuity equation gives
Z h
0
∂ρ ∂
∂ ∂
+ ðρux Þ +
ρuy + ðρuz Þ dz ¼ 0
∂t ∂x
∂y
∂z
(5.64)
The integral
Z h
∂
∂h ∂
½ f ðx, y, zÞdz ¼ f ðx, y, hÞ +
∂x ∂x
0 ∂x
Z h
f ðx, y, zÞdz
(5.65)
0
If the density is assumed to be the mean density of the fluid across the film, then the ux
term in Eq. (5.64) is
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Mechanical Design Engineering Handbook
Z h
∂
∂h ∂
ðρux Þdz ¼ ðρux Þz¼h +
ρux dz
∂x ∂x 0
0 ∂x
Z h
∂h ∂
¼ ρux, b +
ρux dz
∂x ∂x 0
Z h
(5.66)
Similarly for the uy term in Eq. (5.64).
Z h
∂ ∂h ∂
ρuy dz ¼ ρuy, b +
∂y ∂y
0 ∂y
Z h
ρuy dz
(5.67)
0
The uz term can be integrated directly giving
Z h
∂
ðρuz Þdz ¼ ρðuz, b uz, a Þ
∂z
0
(5.68)
The integral form of the continuity equation (Eq. 5.64) on substitution of Eqs (5.66)–
(5.68), can be stated as
Z h
∂ρ
∂h ∂
ρ
ux dz
h ρux, b +
∂t
∂x ∂x
Z h0 ∂h ∂
ρuy, b +
ρ
uy dz + ρðuz, b uz, a Þ
∂y ∂y
0
¼0
(5.69)
The integrals in Eq. (5.69) represent the volumetric flow rates per unit width. Substitution of the values for these from Eqs (5.62), (5.63) gives the general Reynolds
equation.
∂
ρh3 ∂p
∂
ρh3 ∂p
∂ ρhðux, a + ux, b Þ
∂ ρh uy, a + uy, b
+
+
+
+
12μ ∂x
12μ ∂y
∂x
∂y
∂x
2
∂y
2
ρðuz, b uz, a Þ ρux, b
∂h
∂h
∂p
ρuy, b + h ¼ 0
∂x
∂y
∂t
(5.70)
The first two terms of Eq. (5.70) are the Poiseuille terms and describe the net flow
rates due to pressure gradients within the lubricated area. The third and fourth terms
are the Couette terms and describe the net entrained flow rates due to surface velocities. The fifth to the seventh terms are due to a squeezing motion. The eighth term
describes the net flow rate due to local expansion as a result of density changes with
time. Eq. (5.70) is repeated below with the identification of the various terms
emphasised.
Journal bearings
195
∂
ρh3 ∂p
∂
ρh3 ∂p
∂ ρhðux, a + ux, b Þ
∂ ρh uy, a + uy, b
+
+
+
12μ ∂x
12μ ∂y
∂x
∂y
∂x
2
∂y
2
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Couette
Poiseuille
∂h
∂h
+ρðuz, b uz, a Þ ρux, b ρuy, b +
∂x
∂y
|fflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflffl{zfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflffl}
Squeeze
∂p
h
∂t
|{z}
¼0
Net flow due to local expansion
For tangential only motion, where
uz, b ¼ ux, b
∂h
∂h
+ uy, b
∂x
∂y
(5.71)
and uz,a ¼ 0, Eq. (5.70) reduces to
∂ ρh3 ∂p
∂ ρh3 ∂p
∂ðρhÞ
∂ðρhÞ
+
¼ 12ux
+ 12uy
∂x μ ∂x
∂y μ ∂y
∂x
∂y
(5.72)
where
ux ¼
ux, a + ux, b
¼ constant
2
(5.73)
uy ¼
uy, a + uy, b
¼ constant
2
(5.74)
For hydrodynamic lubrication, the fluid properties do not vary significantly through
the bearing and can be considered constant. In addition for hydrodynamic lubrication,
the motion is pure sliding and uy ¼ 0.
The Reynolds equation can therefore be simplified to
∂
∂
∂h
3 ∂p
3 ∂p
h
+
h
¼ 12ux μo
∂x
∂x
∂y
∂y
∂x
(5.75)
In some lubrication applications, side leakage can be neglected and Eq. (5.72) can
be restated as
∂ ρh3 ∂p
∂ðρhÞ
¼ 12ux
∂x μ ∂x
∂x
(5.76)
Eq. (5.76) can be integrated giving
1 dp 12ux
A
¼ 2 + 3
h
μ dx
ρh
(5.77)
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Mechanical Design Engineering Handbook
With boundary conditions dp/dx ¼ 0 when x ¼ xm, ρ ¼ ρm, h ¼ hm gives
A ¼ 12ux ρm hm . The subscript m refers to the condition for which dp/dx ¼ 0 such
as the point of maximum pressure. Substituting for A in Eq. (5.77) gives
dp
ρh ρm hm
¼ 12ux μ
ρh3
dx
(5.78)
If the density can be considered constant then Eq. (5.78) becomes
dp
h hm
¼ 12ux μ 3
dx
h
(5.79)
For the case of a gas lubricated bearing the density, using the ideal gas law, is given by
ρ¼
p
RT
(5.80)
From Eq. (5.72), and taking the viscosity as constant
∂
∂
∂ðρhÞ
3 ∂p
3 ∂p
ρh
+
ρh
¼ 12ux μo
∂x
∂x
∂y
∂y
∂x
(5.81)
Expressing Eq. (5.72) in cylindrical coordinates
∂ rρh3 ∂p
1 ∂ ρh3 ∂p
∂ðρrhÞ
∂ðρhÞ
+
¼ 12 ur
+ uϕ
μ ∂r
∂r
r ∂ϕ μ ∂ϕ
∂r
∂ϕ
(5.82)
where
ur ¼
ur, a + ur, b
2
(5.83)
uϕ ¼
uϕ, a + uϕ, b
2
(5.84)
If the viscosity and density can be assumed to be constant then Eq. (5.82) becomes
∂
∂p
1 ∂
∂p
∂ðrhÞ
∂h
rh3
+
h3
¼ 12μo ur
+ uϕ
∂r
∂r
r ∂ϕ
∂ϕ
∂r
∂ϕ
(5.85)
Eq. (5.85) applies to a thrust bearing where the fluid film is in the z direction and the
bearing dimensions are in the r, ϕ directions.
For turbulent flow, the Reynolds equation is
∂ h3 ∂p
∂ h3 ∂p
ux ∂h
+
¼
2 ∂x
∂x μkx ∂x
∂y μky ∂y
(5.86)
Journal bearings
197
where (from Constantinescu (1962))
ρΩrh 0:102
kx ¼ 12 + 0:026
μ
(5.87)
ρΩrh 0:091
μ
(5.88)
ky ¼ 12 + 0:0198
For an infinitely wide journal bearing the pressure in the axial direction can
be assumed to be constant. This approach is valid for a length to diameter ratio of
L/D > 2. The integrated form of the Reynolds Equation (Eq. 5.79) on substitution
for ux ¼ ðua + ub Þ=2 ¼ Ωa=2, as ub ¼ 0 for a stationary bearing surface, gives
dp 6μo aΩðh hm Þ
¼
dx
h3
(5.89)
dx ¼ adϕ
(5.90)
dp 6μo a2 Ωðh hm Þ
¼
dϕ
h3
(5.91)
1
cos α ¼ ½h + a + e cos ðπ ϕÞ
b
(5.92)
h ¼ bcos α a + e cos ϕ
(5.93)
Hence, using trigonometric relationships,
0:5
e 2
2
h¼b 1
sin ϕ
a + ecos ϕ
b
(5.94)
Expanding the series
1
e 2
b
0:5
sin 2 ϕ
¼1
1 e 2 2
1 e 4 4
sin ϕ sin ϕ …
2 b
8 b
(5.95)
Hence
1 e 2 2
1 e 4 4
h¼b 1
sin ϕ sin ϕ … a + e cos ϕ
2 b
8 b
(5.96)
ba¼c
(5.97)
h ¼ c + e cos ϕ 1 e
1 e 3 4
sin ϕ …
sin 2 ϕ 2 b
8 b
(5.98)
198
Mechanical Design Engineering Handbook
The ratio e/b is typically of the order of 1 103. As a result terms in Eq. (5.98) with
this ratio can readily be neglected, giving
e
h ¼ c + e cos ϕ ¼ c 1 + cos ϕ ¼ cð1 + ε cos ϕÞ
c
(5.99)
where ε is the eccentricity ratio defined by
ε¼
e
c
(5.100)
Substitution of Eq. (5.99) in Eq. (5.91) gives
"
#
a 2
dp
1
hm
¼ 6μΩ
dϕ
c
ð1 + ε cos ϕÞ2 cð1 + ε cos ϕÞ3
(5.101)
Integration of Eq. (5.101) gives the following expression for the pressure distribution.
a 2 Z
p ¼ 6μΩ
c
"
1
ð1 + εcos ϕÞ2
hm
cð1 + ε cos ϕÞ3
#
dϕ + A
(5.102)
where A is a constant of integration.
For a journal bearing,
x ¼ aϕ
(5.103)
and
u¼
ux, a aΩ
¼
2
2
(5.104)
From Eq. (5.75),
∂
∂h
3 ∂p
2 ∂
3 ∂p
h
+a
h
¼ 6μo Ωa2
∂ϕ
∂ϕ
∂y
∂y
∂ϕ
(5.105)
Substituting for the film thickness gives
∂
∂p
∂2 p
h3
+ a2 h3 2 ¼ 6μo Ωa2 esin ϕ
∂ϕ
∂ϕ
∂y
(5.106)
The Reynolds equation can be solved by approximate mathematical methods or
numerically. Once the pressure distribution has been established the journal performance can be determined in terms of the bearing load capacity, frictional losses, lubricant flow requirements and the lubricant temperature rise.
Journal bearings
5.4.2
199
Design charts for full-film hydrodynamic bearings
If the designer wishes to avoid the direct solution of Eq. (5.3), use can be made of a
series of design charts. These were originally produced by Raimondi and Boyd
(1958a, b, c) who used an iterative technique to solve the Reynolds equations. These
charts give the film thickness, coefficient of friction, lubricant flow, lubricant side
flow ratio, minimum film thickness location, maximum pressure ratio, maximum
pressure ratio position and film termination angle versus the Sommerfield number
(Figs 5.18–5.25).
The Sommerfield number, which is also known as the bearing characteristic number, is defined in Eq. (5.107). It is used as it encapsulates all the parameters usually
defined by the designer. Great care needs to be taken with units in the design of journal
bearings. Many design charts have been produced using English units (psi, reyn, Btu
etc.). As long as a consistent set of units is maintained, use of the charts will yield
sensible results. In particular note the use of revolutions per second in the definition
of speed in the Sommerfield number.
S¼
r 2 μN
c
s
P
(5.107)
where:
S is the bearing characteristic number;
r is the journal radius, (m);
c is the radial clearance (m);
μ is the absolute viscosity (Pa s);
Ns is the journal speed (revolutions per second, rps);
P ¼ W/LD is the load per unit of projected bearing area (N/m2);
W is the load on the bearing (N);
D is the journal diameter (m);
L is the journal bearing length (m).
Consider the journal shown in Fig. 5.26. As the journal rotates it will pump lubricant in
a clockwise direction. The lubricant is pumped into a wedge-shaped space and the
journal is forced over to the opposite side. The angular position where the lubricant
film is at its minimum thickness ho is called the attitude angle, ϕ. The centre of the
journal is displaced from the centre of the bearing by a distance e called the eccentricity. The ratio of the eccentricity e to the radial clearance c is called the eccentricity
ratio (ε ¼ e/c). The relationship between the film thickness, the clearance, eccentricity
and eccentricity ratio are defined in Eqs (5.108)–(5.110).
ho ¼ c e
(5.108)
e
c
(5.109)
h0
¼1ε
c
(5.110)
ε¼
200
1
0.9
L/D = inf
Minimum film
thickness variable ho / c
0.8
0.7
0.6
L/D = 1
L/D = 0.5
L/D = 0.25
Max W
0.5
Min f
0.4
0.3
0.1
0
0.001
0.01
0.1
1
2
Sommerfield number, S = (a/c) µNs /P
Fig. 5.18 Chart for the minimum film thickness variable, (ho/c) versus the Sommerfield number.
Data from Raimondi and Boyd (1958c).
10
Mechanical Design Engineering Handbook
0.2
Journal bearings
Coefficient of friction variable, (r/c)f
1000
100
10
1
0.1
0.001
L/D = inf
L/D = 1
L/D = 0.5
L/D = 0.25
0.01
0.1
1
10
Sommerfield number, S = (r/c) mNs /P
2
201
Fig. 5.19 Chart for determining the coefficient of friction variable, (r/c)f.
Data from Raimondi, A.A., Boyd, J., 1958c. A solution for the finite journal bearing and its application to analysis and design: III. ASLE
Trans. 1, 194–209.
202
7
Flow variable, Q/rcNsL
6
5
4
3
2
1
0
0.001
0.01
0.1
1
10
100
Sommerfield number, S = (r/c) mNs /P
2
Fig. 5.20 Chart for determining the flow variable, Q/(acNsL).
Data from Raimondi, A.A., Boyd, J., 1958c. A solution for the finite journal bearing and its application to analysis and design: III. ASLE Trans. 1,
194–209.
Mechanical Design Engineering Handbook
L/D = inf
L/D = 1
L/D = 0.5
L/D = 0.25
Journal bearings
1.2
Flow ratio Qs/Q
1
0.8
0.6
0.4
0.2
0
0.001
L/D = 1
L/D = 0.5
L/D = 0.25
0.01
0.1
1
10
Sommerfield number, S = (r/c) mNs /P
2
203
Fig. 5.21 Chart for determining the ratio of side flow, Qs, to total flow, Q.
Data from Raimondi, A.A., Boyd, J., 1958c. A solution for the finite journal bearing and its application to analysis and design: III. ASLE
Trans. 1, 194–209.
Position of minimum film thickness f(degree)
100
90
80
70
60
50
40
30
L/D = inf
L/D = 1
L/D = 0.5
L/D = 0.25
20
10
0
0.001
0.01
0.1
1
10
Sommerfield number, S = (r/c)2mNs /P
Fig. 5.22 Chart for determining the position of the minimum film thickness ϕ.
Data from Raimondi, A.A., Boyd, J., 1958c. A solution for the finite journal bearing and its
application to analysis and design: III. ASLE Trans. 1, 194–209.
0.9
Maximum film pressure ratio P/Pmax
0.8
L/D = inf
L/D = 1
L/D = 0.5
L/D = 0.25
0.7
0.6
0.5
0.4
0.3
0.2
0.1
0
0.001
0.01
0.1
1
10
Sommerfield number, S = (r/c)2mNs /P
Fig. 5.23 Chart for determining the maximum film pressure ratio, p/pmax.
Data from Raimondi, A.A., Boyd, J., 1958c. A solution for the finite journal bearing and its
application to analysis and design: III. ASLE Trans. 1, 194–209.
Location of maximum film pressure (degree)
25
L/D = inf
L/D = 1
L/D = 0.5
L/D = 0.25
20
15
10
5
0
0.001
0.01
0.1
1
10
Sommerfield number, S = (r/c)2mNs /P
Fig. 5.24 Chart for determining the position of maximum pressure, φpmax.
Data from Raimondi, A.A., Boyd, J., 1958c. A solution for the finite journal bearing and its
application to analysis and design: III. ASLE Trans. 1, 194–209.
160
Terminating position of film (degree)
140
120
100
80
60
40
L/D = inf
L/D = 1
L/D = 0.5
L/D = 0.25
20
0
0.001
0.01
0.1
1
10
Sommerfield number, S = (r/c)2mNs /P
Fig. 5.25 Chart for determining the film termination angle, φpo.
Data from Raimondi, A.A., Boyd, J., 1958c. A solution for the finite journal bearing and its
application to analysis and design: III. ASLE Trans. 1, 194–209.
206
Mechanical Design Engineering Handbook
Start
Speeding up
Steady speed
e
ho
Journal
Bearing
Centre line
shift
Centre line
shift
Fig. 5.26 Full film hydrodynamic bearing motion from start up.
One of the assumptions made in the analysis of Raimondi and Boyd is that the viscosity of the lubricant is constant as it passes through the bearing. However, work is done
on the lubricant in the bearing and the temperature of the lubricant leaving the bearing
zone will be higher than the entrance value. Fig. 5.7 shows the variation of viscosity
with temperature for some of the SAE and ISO defined lubricants and it can be seen
the value of the viscosity for a particular lubricant is highly dependent on the temperature. Some of the lubricant entrained into the bearing film emerges as side flow which
carries away some of the heat. The remainder flows through the load-bearing zone and
carries away the remainder of the heat generated. The temperature used for determining the viscosity can be taken as the average of the inlet and exit lubricant temperatures as given by Eq. (5.111).
Tav ¼ T1 +
T2 T1
ΔT
¼ T1 +
2
2
(5.111)
where T1 is the temperature of the lubricant supply and T2 is the temperature of the
lubricant leaving the bearing. Generally petroleum lubrication oils should be limited
to a maximum temperature of approximately 70°C to prevent excessive oxidation.
One of the parameters that needs to be determined is the bearing lubricant exit temperature T2. This is a trial and error or iterative process. A value of the temperature
rise, ΔT, is guessed, the viscosity for a standard oil, corresponding to this value, determined and the analysis performed. If the temperature rise calculated by the analysis
does not correspond closely to the guessed value the process should be repeated using
an updated value for the temperature rise until the two match. If the temperature rise is
unacceptable it may be necessary to try a different lubricant or modify the bearing
configuration.
Given the length, the journal radius, the radial clearance, the lubricant type and its
supply temperature the steps for determining the various bearing operating parameters
are given in Table 5.4.
The temperature rise of the lubricant through the bearing can be determined from
Journal bearings
207
Table 5.4 Full-film hydrodynamic bearing design guideline
Start Selection
Determine speed and load
Determine the speed of the journal and the
load to be supported.
No
OK
Determine bearing L and D
No
OK
Determine the radial clearance, c
No
OK
Select a lubricant
No
OK
If L and D not already determined, set the
proportions of the bearing so that the load
capacity, P = W/LD, is somewhere
between 0.34 MN/m2 for light machinery to
13.4 MN/m2 for heavy machinery
Determine a value for the radial clearance,
c, of the bearing using the data presented
in Figure 5.14.
If not already specified select a lubricant.
Lubricant oil selection is a function of speed
or compatibility with other lubricant
requirements. Generally as the design
speed rises, oils with a lower viscosity
should be selected.
Estimate a value for the temperature rise
DT across the bearing. The value taken for
Estimate DT
No
OK
Calculate Tav
No
OK
Calculate the Sommerfield number
Determine
the
average
lubricant
temperature Tav=T1+ DT/2 and find the
corresponding value for the viscosity for the
chosen lubricant.
Calculate the Sommerfield number, S =
(r/c)2 µNs/P and the length to diameter ratio.
No
OK
Determine (r/c)f, Q/(rcNsL), (Qs/Q)
2
the initial estimate is relatively
unimportant. As a guide, a value of DT = 10
°C is generally a good starting guess. This
value can be increased for high speed
bearings and for low bearing clearances
No
OK
Use the charts (Figs. 5.19, 5.20 and 5.21)
to determine values for the coefficient of
friction variable, the total lubricant flow
variable, and the ratio of the side flow to the
total lubricant flow with the values for the
Sommerfield number and the L/D ratio.
1
Continued
208
Mechanical Design Engineering Handbook
Table 5.4 Continued
1
2
Calculate T
Calculate the temperature rise of the
lubricant through the bearing using
T
OK
Does T match previous value
No
OK
Check operation in optimal zone
OK
Calculate Q, Qs, f
OK
Calculate ho,p/pmax,fpmax,fpo
OK
Calculate the torque
OK
Calculate the power lost
OK
Specify surface roughness
OK
Design lubricant system
OK
Complete
8.30 10 6 P
r /c f
1 21 Qs / Q Q / rcNsL
If the calculated value does not match the
estimated value for T to within say 1°C
repeat the procedure from the calculation
of Tav using the updated value of the
temperature rise to determine the average
lubricant temperature.
Check the values for S and L/D ratio give a
design that is in the optimal operating
region for minimal friction and maximum
load capability on the chart for the minimum
film thickness variable (Fig. 5.18). If the
operating Sommerfield number and L/D
ratio combination do not fall within this zone,
then it is likely that the bearing design can
be improved by altering values for c, L, D,
the lubricant and the operating temperature.
If the temperature rise across the bearing
has converged, calculate values for the
total lubricant flow rate, Q, the side flow rate
Qs and the coefficient of friction f.
Figs. 5.18 to 5.25 can be used to determine
values for the maximum film thickness, the
maximum film pressure ratio, the location
of the maximum film pressure and its
terminating angular location as required.
The torque required to overcome friction in
the bearing can be calculated by
Torque = fWr
The power lost in the bearing is given by
Power = w Torque = 2pNs Torque
Specify the surface roughness for the
journal and bearing surfaces. A ground
journal with an arithmetic average surface
roughness of 0.4 to 0.8
m is
recommended for bearings of good quality.
For high precision equipment surfaces can
be lapped or polished to give a surface
roughness of 0.2 to 0.4 m. The specified
roughness should be less than the
minimum film thickness.
Design the recirculation and
system for the bearing lubricant.
sealing
Journal bearings
ΔT ¼
209
8:30 106 P
ðr=cÞf
1
Q=rcN s L
1 ðQs =QÞ
2
(5.112)
The torque required to overcome friction in the bearing can be calculated by
Torque ¼ fWr
(5.113)
The power lost in the bearing is given by
Power ¼ ω Torque ¼ 2πNs Torque
(5.114)
where ω is the angular velocity (rad/s).
For values of the length to diameter (L/D) ratio other than those shown in the charts
given in Figs 5.18–5.25, values for the various parameters can be found by interpolation using Eq. (5.115)(Raimondi and Boyd, 1958b).
3
1
L
L
L
1
1
2
1
4
y
+
∞
7
1 6
D D
6 8
7
D
y¼
5
34 1
L
L
1
L
L
1
L
L
ðL=DÞ
12
14
y1 1 14
y1=2 +
1
12
y1=4
3
D
D
4
D
D
24
D
D
2
(5.115)
Here y is the desired variable and y∞, y1, y1/2 and y1/4 are the variable at L/D ¼ ∞, 1,
1/2 and 1/4 respectively.
Ø50
Ø50.08
Example 5.3. A full journal bearing has a nominal diameter of 50.0 mm and a bearing
length of 25.0 mm (Fig. 5.27). The bearing supports a load of 3000 N, and the journal
design speed is 3000 rpm. The radial clearance has been specified as 0.04 mm. An SAE
10 oil has been chosen and the lubricant supply temperature is 50°C.
Find the temperature rise of the lubricant, the lubricant flow rate, the minimum film
thickness, the torque required to overcome friction and the heat generated in the
bearing.
3000 N
25.0
Fig. 5.27 Bearing design example.
210
Mechanical Design Engineering Handbook
Solution
The primary data are D ¼ 50.0 mm, L ¼ 25 mm, W ¼ 3000 N, N ¼ 3000 rpm,
c ¼ 0.04 mm, SAE 10, T1 ¼ 50°C.
Guess a value for the lubricant temperature rise ΔT across the bearing to be say
ΔT ¼ 20°C.
Tav ¼ T1 +
ΔT
20
¼ 50 + ¼ 60° C
2
2
From Fig. 5.7 for SAE 10 at 60°C, μ ¼ 0.014 Pa s.
The mean radius is given by
rm ¼
a + b 25 + 25:04
¼
¼ 25:02 mm
2
2
b a ¼ 0:04mm
Assuming a density of 900 kg/m3,
ν ¼ μ=ρ ¼ 0:014=900 ¼ 1:55 105 m2 =s
So the Taylor number using this approximate value for the kinematic viscosity is
given by
Tam ¼
Ωrm0:5 ðb aÞ1:5
¼ 0:81
ν
As this is much less than the critical Taylor number, 41.19, laminar flow can be
assumed.
Ns ¼ 3000=60 ¼ 50rps, L=D ¼ 25=50 ¼ 0:5:
P¼
S¼
W
3000
¼
¼ 2:4 106 N=m2
LD 0:025 0:05
r 2 μN
c
25 103
¼
P
0:04 103
s
2
0:014 50
¼ 0:1139
2:4 106
From Fig. 5.19 with S ¼ 0.1139 and L/D ¼ 0.5, (r/c)f ¼ 3.8.
From Fig. 5.20 with S ¼ 0.1139 and L/D ¼ 0.5, Q/(rcNsL) ¼ 5.34.
From Fig. 5.21 with S ¼ 0.1139 and L/D ¼ 0.5, Qs/Q ¼ 0.852.
The value of the temperature rise of the lubricant can now be calculated using
Eq. (5.112):
r
f
8:3 106 P
8:3 106 2:4 106 3:8
c
¼ 24:70° C
¼
ΔT ¼
Qs
Q
1 ð0:5 0:852Þ
5:34
1 0:5
Q
rcN s L
Journal bearings
211
As the value calculated for the lubricant temperature rise is significantly different to
the estimated value, it is necessary to repeat the above calculation but using the new
improved estimate for determining the average lubricant temperature.
°
Using ΔT ¼ 24.70°C to calculate Tav gives: Tav ¼ 50 + 24:70
2 ¼ 62:35 C.
Repeating the procedure using the new value for Tav gives:
μ ¼ 0:0136Pas,
S ¼ 0:1107,
from Fig. 5.19 with S ¼ 0.1107 and L/D ¼ 0.5, (r/c)f ¼ 3.7,
from Fig. 5.20 with S ¼ 0.1107 and L/D ¼ 0.5, Q/(rcNsL) ¼ 5.35,
from Fig. 5.21 with S ¼ 0.1107 and L/D ¼ 0.5,Qs/Q ¼ 0.856,
ΔT ¼
8:3 106 2:4 106 3:7
¼ 24:08° C,
1 ð0:5 0:856Þ
5:35
Tav ¼ 50 +
24:08
¼ 62:04° C:
2
This value for Tav is close to the previous calculated value suggesting that the solution
has converged. For Tav ¼ 62.04°C, μ ¼ 0.0136 Pa s and S ¼ 0.1107.
The other parameters can now be found.
Q ¼ rcN s L 5:35 ¼ 25 0:04 50 25 5:35 ¼ 6688mm3 =s:
From Fig. 5.18ho/c ¼ 0.22. ho ¼ 0.0088 mm.
f ¼ 3:7 ðc=r Þ ¼ 3:7 ð0:04=25Þ ¼ 0:00592:
The torque is given by Torque ¼ fWr ¼ 0.00592 3000 0.025 ¼ 0.444 N m.
The
power
dissipated
in
the
bearing
is
given
Power ¼ 2π Torque Ns ¼ 139.5 W.
by
Example 5.4. A full journal bearing is required for a 30 mm diameter shaft, rotating
at 1200 rpm supporting a load of 1500 N (Fig. 5.28). From previous experience, the
lubrication system available is capable of delivering the lubricant to the bearing at
40°C. Select an appropriate radial clearance, length and lubricant for the bearing
and determine the temperature rise of the lubricant, the total lubricant flow rate
required and the power absorbed in the bearing. Check that your chosen design
operates within the optimum zone indicated by the dotted lines on the minimum film
thickness ratio chart (Fig. 5.18).
Solution
The primary design information is: D ¼ 30.0 mm, W ¼ 1500 N, N ¼ 1200 rpm, SAE
10, T1 ¼ 40°C.
212
Mechanical Design Engineering Handbook
Ø30
1500 N
L
Fig. 5.28 Bearing design example.
From Fig. 5.14 for a speed of 1200 rpm and a nominal journal diameter of 30 mm, a
suitable value for the diametral clearance is 0.05 mm. The radial clearance is therefore
0.025 mm. The next step is to set the length of the bearing. Typical values for L/D
ratios are between 0.5 and 1.5. Here a value for the ratio is arbitrarily set as L/
D ¼ 1 (i.e. L ¼ 30 mm).
Given the speed of the bearing, 1200 rpm, as an initial proposal an SAE 10 lubricant
is selected.
The procedure for determining the average lubricant temperature is the same as for
the previous example. As a first estimate the temperature rise of the lubricant is taken
as 10°C and the results of the iterative procedure are given in Table 5.5.
P¼
S¼
W
1500
¼
¼ 1:667 106 N=m2 :
LD 0:03 0:03
r 2 μN
c
Tav ¼ T1 +
P
s
¼
15 103
0:025 103
2
μ20
¼ 4:319μ:
1:667 106
ΔT
10
¼ 40 + ¼ 45° C:
2
2
The mean radius is given by
a + b 15 + 15:025
¼
¼ 15:01mm
rm ¼
2
2
b a ¼ 0:025mm
Table 5.5 Tabular data for Example 5.4
Temperature rise, ΔT (°C)
Average lubricant temperature Tav (°C)
Average lubricant viscosity μ(Pa s)
Sommerfield number, S
Coefficient of friction variable (r/c)f
Flow variable Q/rcNsL
Side flow to total flow ratio Qs/Q
10
45
0.028
0.1209
3.3
4.35
0.68
15.90
47.95
0.023
0.09934
2.8
4.4
0.72
13.76
46.88
0.024
0.1037
2.8
4.4
0.72
13.76
46.88
Converged
Journal bearings
213
An estimate for the kinematic viscosity can be found, taking the average temperature
of the oil as 45°C and assuming a density of 900 kg/m3. At 45°C, the viscosity for SAE
10 is approximately 0.028 Pa s
ν ¼ μ=ρ ¼ 0:028=900 ¼ 3:11 105 m2 =s
So the Taylor number using this approximate value for the kinematic viscosity is
given by
Tam ¼
Ωrm0:5 ðb aÞ1:5
¼ 0:062
ν
As this is much less than the critical Taylor number, 41.19, laminar flow can be
assumed.
The total lubricant flow rate required is given by Q ¼ rcNsL
4.4 ¼ 15 0.025 20 30 4.4 ¼ 990 mm3/s.
With S ¼ 0.1037 and L/D ¼ 1 the design selected is within the optimum operating
zone for minimum friction and optimum load capacity indicated in Fig. 5.18.
The friction factor, f ¼ 2.8 (c/r) ¼ 2.8 0.025/15 ¼ 0.004667.
Torque ¼ fWr ¼ 0.004667 1500 0.015 ¼ 0.105 N m.
Power ¼ 2πNs Torque ¼ 13.19 W.
Example 5.5. A full journal bearing has a nominal diameter of 25.0 mm and bearing
length 25.0 mm. The bearing supports a load of 1450 N and the journal design speed is
2500 rpm. Two radial clearances are proposed (i) 0.025 mm and (ii) 0.015 mm. An
SAE 10 oil has been chosen and the lubricant supply temperature is 50°C. Using
the standard Raimondi and Boyd design charts, find the temperature rise of the
lubricant, the lubricant flow rate, the torque required to overcome friction and the
heat generated in the bearing for both of the proposed clearances. Determine which
clearance, if any, should be selected for the bearing.
Solution
D ¼ 25 mm, L ¼ 25 mm, W ¼ 1450 N, N ¼ 2500 rpm, Ns ¼ 2500/60 rps ¼ 41.67 rps,
SAE 10, T1 ¼ 50°C, L/D ¼ 1.
P¼
W
1450
¼
¼ 2:32 106 Nm2 :
LD 0:025 0:025
(i) c ¼ 0.025 mm.
S¼
r 2 μN
c
P
s
¼
12:5 103
0:025 103
The mean radius is given by
2
μ41:67
¼ 4:49μ:
2:32 106
214
Mechanical Design Engineering Handbook
rm ¼
a + b 12:5 + 12:525
¼
¼ 12:51mm
2
2
b a ¼ 0:025mm
An estimate for the kinematic viscosity can be found, taking the average temperature
of the oil as 55°C and assuming a density of 900 kg/m3. At 55°C, the viscosity for SAE
10 is approximately 0.017 Pa s
ν ¼ μ=ρ ¼ 0:017=900 ¼ 1:89 105 m2 =s
So the Taylor number using this approximate value for the kinematic viscosity is given by
Tam ¼
Ωrm0:5 ðb aÞ1:5
¼ 0:19
ν
As this is much less than the critical Taylor number, 41.19, laminar flow can be
assumed. Reducing the clearance results in a smaller Taylor number so laminar flow
can also be assumed for the smaller clearance.
The solution for the iterative procedure using the Raymondi and Boyd charts
(Figs 5.19–5.21) to find the average temperature rise of the lubricant across the bearing is given in tabular form below. The first estimate for the temperature rise is taken
as ΔT ¼ 10°C.
Temperature rise, ΔT (°C)
Average lubricant temperature Tav (°C)
Average lubricant viscosity μ(Pa s)
Sommerfield number, S
Coefficient of friction variable (r/c)f
Flow variable Q/rcNsL
Side flow to total flow ratio Qs/Q
10
55
0.017
0.07633
2.3
4.45
0.766
16.13
58.07
0.015
0.06735
2.2
4.52
0.78
15.36
57.68
Converged
ΔT 15:4° C:
Q ¼ rcN s L 4:52 ¼ 12:5 0:025 41:67 25 4:52 ¼ 1471mm3 =s:
f ¼ 2:2 ðc=r Þ ¼ 2:2 0:025=12:5 ¼ 0:0044:
T ¼ fWr ¼ 0:0044 1450 0:0125 ¼ 0:07975Nm:
Heat generated ¼ 2πNs Torque ¼ 20:88W:
With S ¼ 0.06735 and L/D ¼ 1 the design selected is outside the optimum operating
zone for minimum friction and optimum load capacity indicated in Fig. 5.18.
Journal bearings
215
(ii) c ¼ 0.015 mm.
S¼
r 2 μN
c
P
s
¼
12:5 103
0:015 103
2
μ41:67
¼ 12:47μ:
2:32 106
Assuming ΔT ¼ 10°C.
Temperature rise, ΔT (°C)
Average lubricant temperature Tav (°C)
Average lubricant viscosity μ(Pa s)
Sommerfield number, S
Coefficient of friction variable (r/c)f
Flow variable Q/rcNsL
Side flow to total flow ratio Qs/Q
10
55
0.017
0.2120
4.8
4.08
0.55
27.01
63.51
0.0125
0.1559
3.8
4.22
0.622
25.17
62.58
0.013
0.1621
3.9
4.21
0.62
25.85
62.93
Converged
ΔT 25:9° C:
Q ¼ rcN s L 4:21 ¼ 12:5 0:015 41:67 25 4:21 ¼ 822:3mm3 =s:
f ¼ 3:9 ðc=r Þ ¼ 3:9 0:015=12:5 ¼ 0:00468:
T ¼ fWr ¼ 0:00468 1450 0:0125 ¼ 0:08483Nm:
Heat generated ¼ 2πNs Torque ¼ 22:21W:
With S ¼ 0.1621 and L/D ¼ 1 the design selected is inside the optimum operating zone
for minimum friction and optimum load capacity indicated in Fig. 5.18.
Therefore this clearance (c ¼ 0.015 mm) should be selected.
Example 5.6. A full film hydrodynamic journal bearing is required for a shaft
supporting a radial load of 7500 N whilst rotating at 800 rpm. The nominal diameter
of the shaft is 80 mm. Select a suitable lubricant and appropriate values for the radial
clearance and bearing length. Determine the overall temperature rise in the bearing
and the flow rate of lubricant required. Establish that the selected design lies within
the optimum zone for minimum friction and maximum load on the minimum film
thickness ratio chart. The lubricant can be supplied at an inlet temperature of 50°C
using an available cooling system.
Solution
From Fig. 5.14 for N ¼ 800 rpm and D ¼ 80 mm, 2c ¼ 80 μm.
Hence c ¼ 40 μm.
13:4 > P > 0:34 ðMPaÞ
P¼
W
7500
¼
LD 0:08L
216
Mechanical Design Engineering Handbook
Try L/D ¼ 0.5
P¼
W
7500
¼
¼ 2:34MPa
LD 0:08 0:04
Try SAE 20. As Ns + SAE number *
The mean radius is given by
rm ¼
a + b 40 + 40:04
¼
¼ 40:02 mm
2
2
b a ¼ 0:04mm
An estimate for the kinematic viscosity can be found, taking the average temperature
of the oil as 65°C and assuming a density of 900 kg/m3. At 55°C, the viscosity for SAE
20 is approximately 0.017 Pa s
ν ¼ μ=ρ ¼ 0:017=900 ¼ 1:89 105 m2 =s
So the Taylor number using this approximate value for the kinematic viscosity is given
by
Tam ¼
Ωrm0:5 ðb aÞ1:5
¼ 0:22
ν
As this is much less than the critical Taylor number, 41.19, laminar flow can be
assumed.
S¼
r 2 μN
c
40 103
¼
P
0:04 103
2
s
Temperature rise, ΔT (°C)
Average lubricant temperature Tav (°C)
Average lubricant viscosity μ(Pa s)
Sommerfield number, S
Coefficient of friction variable (r/c)f
Flow variable Q/rcNsL
Side flow to total flow ratio Qs/Q
μ13:33
¼ 5:69μ
2:34 106
30
65
0.017
0.0967
3.4
5.4
0.866
21.6
60.8
0.02
0.114
3.7
5.35
0.84
23.2
61.66
0.02
Converged
With S ¼ 0.114 and L/D ¼ 0.5, the design is well within the optimum zone for minimum friction and maximum load.
Q ¼ rcN s L 5:35 ¼ 40 0:04 13:33 40 5:35 ¼ 4565mm3 =s:
Journal bearings
217
Example 5.7. A design proposal for a full film hydrodynamic bearing comprises a
bearing of nominal diameter 30 mm, length 15 mm using ISO VG 22 lubricant supplied at 50°C. The bearing load is 1280 N and the rotational speed is 6000 rpm.
The radial clearance for the bearing is 0.022 mm. Use the viscosity chart given in
Fig. 5.7 and the Raimondi and Boyd design charts given in Figs 5.18–5.21 to determine the lubricant flow rate, lubricant temperature rise and minimum film thickness.
The initial guess for the temperature rise across the bearing can be taken as 30°C.
If in place of ISO VG 22 oil, SAE 10 oil was used in the bearing determine the
resulting temperature rise of the lubricant.
Comment on the suitability of the proposed design and justify the selection of either
ISO VG 22 or SAE 10 as a lubricant.
Solution
Initial data: L ¼ 15 mm, D ¼ 30 mm. L/D ¼ 0.5.
N ¼ 6000rpm, Ns ¼ 100 rps,
c ¼ 0:022mm:
ISO VG 22.
T1 ¼ 60° C:
ΔT ¼ 30° C
W ¼ 1280N
Tav ¼ 50 +
30
¼ 65° C
2
The mean radius is given by
rm ¼
a + b 15 + 15:022
¼
¼ 15:01mm
2
2
b a ¼ 0:022mm
An estimate for the kinematic viscosity can be found, taking the average temperature
of the oil as 65°C and assuming a density of 900 kg/m3. At 55°C, the viscosity for ISO
VG 22 is approximately 0.0074 Pa s
ν ¼ μ=ρ ¼ 0:0074=900 ¼ 8:22 106 m2 =s
So the Taylor number using this approximate value for the kinematic viscosity is
given by
Tam ¼
Ωrm0:5 ðb aÞ1:5
¼ 0:97
ν
218
Mechanical Design Engineering Handbook
As this is much less than the critical Taylor number, 41.19, laminar flow can be assumed.
P¼
W
1280
¼
¼ 2:844 106 Pa
LD 0:015 0:03
Temperature rise, ΔT (°C)
Average lubricant temperature Tav (°C)
Average lubricant viscosity μ(Pa s)
Sommerfield number, S
Coefficient of friction variable (r/c)f
Flow variable Q/rcNsL
Side flow to total flow ratio Qs/Q
30
65
0.0074
0.12
4
5.3
0.842
30.8
65.4
Converged
ΔT ¼ 31° C
Q ¼ rcN s L 5:3 ¼ 15 0:022 100 15 5:3 ¼ 2624mm3 =s:
ho =c ¼ 0:24
ho ¼ 0:24 0:022mm ¼ 5:28 103 mm
ho ¼ 5:3 μm
From the graph for ho/c versus Sommerfield number, the design is in the optimum
zone for maximum load and minimum friction.
For SAE 10:
Temperature rise, ΔT (°C)
Average lubricant temperature Tav (°C)
Average lubricant viscosity μ(Pa s)
Sommerfield number, S
Coefficient of friction variable (r/c)f
Flow variable Q/rcNsL
Side flow to total flow ratio Qs/Q
30
65
0.0116
0.19
5.6
5.1
0.8
43.2
71.6
0.0092
0.15
4.8
5.2
0.82
36.9
68.5
0.0102
0.167
5
5.15
0.812
38.6
69.3
0.01
Converged
ΔT 38:6° C
With S ¼ 0.167 and L/D ¼ 0.5, the design also falls within the optimum zone for maximum load and minimum friction.
Use ISO VG 22 as the lubricant temperature rise is lower unless other components
or considerations dictate the use of SAE 10.
Journal bearings
5.4.3
219
Alternative method for the design of full film hydrodynamic
bearings
An alternative method to using the design charts of Raimondi and Boyd is to use an
approximate method developed by Reason and Narang (1982). They defined an
approximation for the pressure variation in a journal bearing given by:
)
(
3μV L2 2
ε sin θ
z
rc2 4
ð1 + ε cos θÞ3
p¼
2 + ε2 L2 2
1
z
1+
2r 2 4
ð1 + εcos θÞð2 + ε cos θÞ
(4.116)
where:
μ ¼ viscosity (Pa s),
V ¼ surface velocity (m/s),
r ¼ journal radius (m),
c ¼ radial clearance (m),
L ¼ bearing length (m),
z ¼ axial coordinate (m),
ε ¼ eccentricity variable,
θ ¼ circumferential coordinate.
This equation was integrated using an approximate method and the solution given in
terms of two integrals Is and Ic, the sine and cosine load integrals respectively. Both Is
and Ic are functions of the L/D ratio and the eccentricity ratio ε. The values of these
integrals are given in Table 5.6 along with the corresponding Sommerfield number
calculated using Eq. (5.117).
S¼
6π
1
pffiffiffiffiffiffiffiffiffiffiffiffi
Ic2 + Is2
(5.117)
Knowing values for the integrals Is and Ic enables the bearing parameters to be
calculated as shown below.
l
l
l
l
l
Given: W, Ns, L, D, c, the type of lubricant and T1.
Guess a value for the temperature rise of the lubricant ΔT and calculate Tav. Determine
the viscosity at Tav from Fig. 5.7.
W
Calculate the bearing loading pressure P ¼ LD
.
r 2
μNs
Calculate the Sommerfield number S ¼
.
c P
Knowing the length to diameter ratio and the Sommerfield number, search the correct L/D
column in Table 5.6 for a matching Sommerfield number. A matching Sommerfield number
will give the eccentricity ratio reading across the table to the left. Interpolation may be necessary for values of S between the values listed in the table. Once the correct eccentricity has
been found, determine the values for the sine and cosine integrals Is and Ic to correspond with
the value of the Sommerfield number.
220
Table 5.6 Values of the sine and cosine integrals, Is and Ic, and the Sommerfield number S, as functions of the L/D ratio and the eccentricity ratio
ε(Reason and Narang (1982)
L/D 5 0.25
L/D 5 0.5
L/D 5 0.75
L/D 5 1.0
L/D 5 1.5
L/D 5 2
L/D 5∞
0.10
Is ¼ 0.0032
Ic ¼ 0.0004
S ¼ 16.4506
Is ¼ 0.0067
Ic ¼ 0.0017
S ¼ 7.6750
Is ¼ 0.0109
Ic ¼ 0.0043
S ¼ 4.5276
Is ¼ 0.0164
Ic ¼ 0.0089
S ¼ 2.8432
Is ¼ 0.0241
Ic ¼ 0.0174
S ¼ 1.7848
Is ¼ 0.0363
Ic ¼ 0.0338
S ¼ 1.0696
Is ¼ 0.0582
Ic ¼ 0.0703
S ¼ 0.5813
Is ¼ 0.1071
Ic ¼ 0.1732
S ¼ 0.2605
0.0120
0.0014
4.3912
0.0251
0.0062
2.0519
0.0404
0.0153
1.2280
0.0597
0.0312
0.7876
0.0862
0.0591
0.5076
0.1259
0.1105
0.3167
0.1927
0.2161
0.1832
0.3264
0.4797
0.0914
0.0244
0.0028
2.1601
0.0505
0.0118
1.0230
0.0804
0.0289
0.6209
0.1172
0.0579
0.4058
0.1656
0.1065
0.2694
0.2345
0.1917
0.1752
0.3430
0.3549
0.1075
0.5425
0.7283
0.0584
0.0380
0.0041
1.3880
0.0783
0.0174
0.6614
0.1236
0.0419
0.4065
0.1776
0.0825
0.2709
0.2462
0.1484
0.1845
0.3396
0.2590
0.1242
0.4793
0.4612
0.0798
0.7220
0.8997
0.0460
0.0636
0.0063
0.8301
0.1300
0.0259
0.4002
0.2023
0.0615
0.2509
0.2847
0.1183
0.1721
0.3835
0.2065
0.1218
0.5102
0.3474
0.0859
0.6878
0.5916
0.0585
0.9771
1.0941
0.0362
0.0839
0.0076
0.6297
0.1705
0.0312
0.3061
0.2628
0.0733
0.1944
0.3649
0.1391
0.1359
0.4831
0.2391
0.0984
0.6291
0.3949
0.0714
0.8266
0.6586
0.0502
1.1380
1.1891
0.0322
0.1570
0.0100
0.3372
0.3143
0.0408
0.1674
0.4727
0.0946
0.1100
0.6347
0.1763
0.0850
0.8061
0.2962
0.0618
0.9983
0.4766
0.0480
1.2366
0.7717
0.0364
1.5866
1.3467
0.0255
0.20
0.30
0.40
0.50
0.60
0.70
0.80
Mechanical Design Engineering Handbook
ε
0.95
0.99
Is ¼ 0.2761
Ic ¼ 0.6644
S ¼ 0.0737
Is ¼ 0.6429
Ic ¼ 2.1625
S ¼ 0.0235
Is ¼ 3.3140
Ic ¼ 22.0703
S ¼ 0.0024
0.7079
1.4990
0.0320
1.3712
3.9787
0.0126
4.9224
28.5960
0.0018
1.0499
2.0172
0.0233
1.8467
4.8773
0.0102
5.6905
30.8608
0.0017
1.3002
2.3269
0.0199
2.1632
5.3621
0.0092
6.1373
31.9219
0.0016
1.6235
2.6461
0.0171
2.5455
5.8315
0.0083
6.6295
32.8642
0.0016
1.8137
2.7932
0.0159
2.7600
6.0396
0.0080
6.8881
33.2602
0.0016
2.3083
3.0339
0.0139
3.2913
6.3776
0.0074
8.7210
33.5520
0.0015
Journal bearings
0.90
221
222
Mechanical Design Engineering Handbook
Recall that for a linear function, y ¼ f(x), the value for y between the bounding points
(x1, y1) and (x2, y2) for a given value of x can be found by linear interpolation:
x x1
y¼
ðy2 y1 Þ + y 1
x2 x1
l
l
Calculate the minimum film thickness ho ¼ c(1 ε).
Calculate the attitude angle ϕ, (degree)
ϕ ¼ tan
l
(5.118)
1
Is
Ic
(5.119)
Calculate the flow variable Q/(rcNsL), where Q is the lubricant supply.
Q
¼ π ð 1 + εÞ
rcN s L
2
2
33
6
77
6
2
6
6
77
D
1
1
1 6
77
ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi
s
s
ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi
πεη1 6
1
η
tanh
1
6
6
77
L
4
4
D 2 η1 1
D 2 η1 155
+
+
2
2 4
2 4
L
L
(5.120)
where
η1 ¼
l
ð1 + εÞð2 + εÞ
ð2 + ε2 Þ
(5.121)
Calculate the flow through the minimum film thickness Qπ.
2
2
33
6
6
77
2
6
6
77
Qπ
D
1
1
1 6
77
ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi
s
s
ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi
¼ π ð1 εÞ + πεη2 6
1
η
tanh
2
6
77
6
rcN s L
L
4
4
D 2 η2 1
D 2 η2 155
+
+
2
2 4
L 2 4
L
(5.122)
where
η2 ¼
Note that
ð1 εÞð2 εÞ
ð2 + ε2 Þ
(5.123)
Journal bearings
tanh
l
1
223
1
1+x
x ¼ ln
2
1x
(5.124)
The flow ratio can be calculated using Eq. (5.125).
Qs
Qπ
¼1
Q
Q
l
(5.125)
Calculate the friction variable (r/c)f
r
(
π
ε
f
¼ 6πS pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi + Is
2
c
3 ð1 ε Þ 2
l
)
(5.126)
Calculate the temperature rise:
ΔT ¼
8:30 106 P
ðr=cÞf
1
Q=rcN s L
1 ðQs =QÞ
2
Although these equations may appear complex the advantage of the method developed
by Reason and Narang is that it is suitable for conventional programming or utilisation
within a spreadsheet.
Ø20
Ø20.04
Example 5.8. A full journal bearing has a nominal diameter of 20.0 mm and a bearing
length of 20.0 mm (Fig. 5.29). The bearing supports a load of 1000 N, and the journal
design speed is 6000 rpm. The radial clearance has been specified as 0.02 mm. An SAE
20 oil has been chosen to be compatible with requirements elsewhere in the machine
and the lubricant supply temperature is 60°C. Using the approximate method of Reason and Narang find the temperature rise of the lubricant, the total lubricant flow rate,
the minimum film thickness, the torque required to overcome friction and the heat
generated in the bearing.
Solution
The primary design data are D ¼ 20 mm, L ¼ 20 mm, W ¼ 1000 N, N ¼ 6000 rpm,
c ¼ 0.02 mm, SAE 20, T1 ¼ 60°C.
1000 N
20 mm
Fig. 5.29 Bearing design example.
224
Mechanical Design Engineering Handbook
°
15
Guess ΔT ¼ 15°C. Hence: Tav ¼ T1 + ΔT
2 ¼ 60 + 2 ¼ 67:5 C.
For SAE 20, at 67.5°C, from Fig. 5.7, μ ¼ 0.016 Pa s.
The mean radius is given by
rm ¼
a + b 20 + 20:02
¼
¼ 20:01mm
2
2
b a ¼ 0:02mm:
Assuming a density of 900 kg/m3.
ν ¼ μ=ρ ¼ 0:016=900 ¼ 1:78 105 m2 =s
So the Taylor number using this approximate value for the kinematic viscosity is given
by
Tam ¼
Ωrm0:5 ðb aÞ1:5
¼ 0:45
ν
As this is much less than the critical Taylor number, 41.19, laminar flow can be
assumed.
W
1000
¼
¼ 2:5MPa:
LD 0:02 0:02
r 2 μN 0:01 2 0:016 100
s
S¼
¼
¼ 0:16:
c
0:00002
P
2:5 106
P¼
From Table 5.6 for L/D ¼ 1, the nearest bounding values of the Sommerfield number
to S ¼ 0.16, are for ε ¼ 0.5 (S ¼ 0.1845) and ε ¼ 0.6 (S ¼ 0.1242).
Interpolating to find the values of ε, Is and Ic which correspond to S ¼ 0.16:
ε¼
0:16 0:1845
ð0:6 0:5Þ + 0:5 ¼ 0:5406,
0:1242 0:1845
Is ¼
0:16 0:1845
ð0:3396 0:2462Þ + 0:2462 ¼ 0:2841,
0:1242 0:1845
Ic ¼
0:16 0:1845
ð0:2590 0:1484Þ + 0:1484 ¼ 0:1933:
0:1242 0:1845
From Eq. (5.120) with η1 ¼ 1.708,
Q
¼ 4:84 2:901 1 1:708 0:9517 tanh 1 ð0:4759Þ ¼ 4:379:
rcN s L
Journal bearings
225
From Eq. (5.122) with η2 ¼ 0.2925,
Qπ
¼ 1:443 + 0:4968 1 0:2925 1:589 tanh 1 ð0:7943Þ ¼ 1:690:
rcN s L
From Eq. (5.125), Qs/Q ¼
0.6141.
r
From Eq. (5.126), f c ¼ 3:016ð1:245 + 0:07679Þ ¼ 3:987.
8:3 106 P
ðr=cÞf
¼ 27:26° C:
ΔT ¼
1 0:5ðQs =QÞ Q=rcN s L
Repeating the procedure with this value for ΔT gives:
Tav ¼ 73.6°C, hence μ ¼ 0.013 Pa s,
S ¼ 0.13, hence ε ¼ 0.5904,
Q/(rcNsL) ¼ 4.492,
Qπ/(rcNsL) ¼ 1.531,
(r/c)f ¼ 3.418.
Using these values gives ΔT ¼ 23.56°C.
Repeating the procedure with this value for ΔT gives:
Tav ¼ 71.78°C, hence μ ¼ 0.014 Pa s,
S ¼ 0.14, hence ε ¼ 0.5738,
Q/(rcNsL) ¼ 4.454,
Qπ/(rcNsL) ¼ 1.585,
(r/c)f ¼ 3.613.
Using these values gives ΔT ¼ 24.83°C. Hence Tav ¼ 72.4°C. This is close enough to
the previous value, indicating that the correct value for Tav has been determined.
The values for the total lubricant supply, minimum film thickness, coefficient of
friction, torque and the heat generated in the bearing can now be determined:
Q ¼ 1782 mm3/s,
ho ¼ c(1 ε) ¼ 0.009 mm,
f ¼ 0.007226,
Torque ¼ 0.072 N m,
Power ¼ 45.4 W.
Example 5.9. A full journal bearing has a nominal diameter of 80.0 mm, bearing
length 60.0 mm. The bearing supports a load of 7500 N, and the journal design speed
is 800 rpm. The radial clearance has been specified as 0.05 mm. An SAE 30 oil has
been chosen and the lubricant supply temperature is 50°C. Using the approximate
method of Reason and Narang for determining the design and performance of full film
hydrodynamic bearings, find the temperature rise of the lubricant, the lubricant flow
rate, the torque required to overcome friction, the heat generated in the bearing and the
minimum film thickness.
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Mechanical Design Engineering Handbook
Solution
L/D ¼ 0.75, P ¼ 1.562 MN m2.
The mean radius is given by
rm ¼
a + b 40 + 40:05
¼
¼ 40:03mm
2
2
b a ¼ 0:05mm
An estimate for the kinematic viscosity can be found, taking the average temperature
of the oil as 60°C and assuming a density of 900 kg/m3. At 55°C, the viscosity for SAE
30 is approximately 0.032 Pa s
ν ¼ μ=ρ ¼ 0:032=900 ¼ 3:56 105 m2 =s
So the Taylor number using this approximate value for the kinematic viscosity is
given by
Tam ¼
Ωrm0:5 ðb aÞ1:5
¼ 0:17
ν
As this is much less than the critical Taylor number, 41.19, laminar flow can be
assumed.
Temperature rise, ΔT (°C)
Average lubricant temperature Tav (°C)
Average lubricant viscosity μ(Pa s)
Sommerfield number, S
ε
Is
Ic
Flow variable Q/rcNsL
Qπ/rcNsL
Side flow to total flow ratio Qs/Q
Coefficient of friction variable (r/c)f
20
60
0.032
0.1748
0.6006
0.2352
0.1928
4.715
1.440
0.6946
4.547
ΔT 19:2° C:
ho ¼ 0:05 103 ð1 0:6006Þ ¼ 19:97 μm 20:0 μm:
Q ¼ 4:715 40 0:05 ð800=60Þ 60 ¼ 7544mm3 =s:
Torque ¼ fWr ¼ 4:547ð0:05=40Þ 7500 0:04 ¼ 1:705Nm:
Heat generated ¼ 2πNs Torque ¼ 142.8 W.
19.16
59.58
Converged
Journal bearings
5.5
227
Conclusions
Bearings are used to support a load whilst allowing relative motion between two elements of a machine. Typical applications are shafts where the shaft rotates relative to a
housing but bearings can also be used to facilitate linear relative motion for, say,
machine slides. This chapter has introduced the wide range of bearing types and their
selection and the design of boundary lubricated and full-film hydrodynamic journal
bearings. Boundary lubricated bearings represent an important class of bearings where
the combination of low speed and load make it possible for surface-to-surface contact
to occur without significant wear and this type of bearing is widely used in plant
machinery and product design. In full film hydrodynamic bearings, rotation results
in high pressures in the lubricant which act to separate the journal from the bearing
enabling prolonged high speed operation.
References
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Hamrock, B.J., 1994. Fundamentals of Fluid Film Lubrication. McGraw Hill, New York.
Harris, T.A., 2001. Rolling Bearing Analysis, fourth ed. John Wiley & Sons, New York.
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analysis and design: III. ASLE Trans. 1, 194–209.
Reason, B.R., Narang, I.P., 1982. Rapid design and performance evaluation of steady state
journal bearings—a technique amenable to programmable hand calculators. ASLE Trans.
25 (4), 429–444.
228
Mechanical Design Engineering Handbook
Reynolds, O., 1886. On the theory of lubrication and its application to Mr Beauchamp Tower’s
experiments including an experimental determination of the viscosity of olive oil. Philos.
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Websites
At the time of going to press the world-wide-web contained useful information relating to this
chapter at the following sites.
www.abc-bearing.com.
www.arb-bearings.com.
www.bocabearings.com.
www.buntingbearings.com/data.html.
www.championbearings.com.
www.cooperbearings.com.
www.fag.de.
www.ggbearings.com.
www.gtweed.com/docs/ebrochures/PlasticBearingDesignGuide-0901.pdf.
www.hb-bearings.com.
www.ina.com.
www.jskbearings.com.
www.mrcbearings.com.
www.nationalbearings.com.
www.newindo.com.
www.nsk-rhp.com.
www.ntn.ca.
qbcbearings.com/TechInfo/PDF/TechT32_34.pdf.
www.railko.co.uk.
www.rbcbearings.com.
www.revolve.com.
www.rosebearings.com.
www.skf.com.
www.snfa-bearings.co.uk.
www.spn.com.tw.
www.sun-sunshine.com.
www.thordonbearings.com.
www.timken.com.
www.torrington.com.
www.toyobearings.com.
hwww.tribology-abc.com/calculators/e3_11.htm.
Standards
American National Standards Institute. (1990). ANSI/AFBMA Standard 9-1990.Load ratings
and fatigue life for ball bearings. ANSI.
American National Standards Institute. (1990). ANSI/AFBMA Standard 11-1990.Load ratings
and fatigue life for roller bearings. ANSI.
Journal bearings
229
American National Standards Institute. (1995). ANSI/ABMA Standard 7-1995.Shaft and
housing fits for metric radial ball and roller bearings (except tapered roller bearings)
conforming to basic boundary plans. ANSI.
British Standards Institution. (1979). BS 3134-1:1979.Metric tapered roller bearings. Specification for dimensions and tolerances of single row bearings.
British Standards Institution. (1982). BS 292: Part 1: 1982. Rolling bearings: ball bearings,
cylindrical and spherical roller bearings. Specification for dimensions of ball bearings,
cylindrical and spherical roller bearings (metric series).
British Standards Institution. (1984). BS 6560:1984. Glossary of terms for rolling bearings.
British Standards Institution. (1991). BS5512: 1991. Method of calculating dynamic load rating
and rating life of rolling bearings.
British Standards Institution. (1992). BS 3134-4:1992, ISO 10317:1992. Metric tapered roller
bearings. Specification for designation system.
British Standards Institution. (1992). BS 4480: Part 1: 1992. Plain bearings metric series.
Sintered bushes. Dimensions and tolerances.
British Standards Institution. (1995). BS ISO 12128:1995. Plain bearings. Lubrication holes,
grooves and pockets. Dimensions, types, designation and their application to bearing
bushes.
British Standards Institution. (1998). BS ISO 15:1998. Rolling bearings. Radial bearings.
Boundary dimensions, general plan.
British Standards Institution. (1999). BS ISO 113:1999. Rolling bearings. Plummer block housings. Boundary dimensions.
British Standards Institution. (1999). BS ISO 3547-1:1999.Plain bearings. Wrapped bushes.
Dimensions.
British Standards Institution. (1999). BS ISO 3547-3:1999.Plain bearings. Wrapped bushes.
Lubrication holes, lubrication grooves and lubrication indentations.
British Standards Institution. (1999). BS ISO/TS 16799:1999. Rolling bearings. Dynamic
load ratings and rating life. Discontinuities in the calculating of basic dynamic load
ratings.
British Standards Institution. (2001). BS ISO 12128:2001. Plain bearings. Lubrication holes,
grooves and pockets. Dimensions, types, designation and their application to bearing
bushes.
British Standards Institution. (2001). BS ISO 2982-2:2001.Rolling bearings. Accessories. Locknuts and locking devices. Dimensions.
British Standards Institution. (2001). BS ISO 3290:2001. Rolling bearings. Balls. Dimensions
and tolerances.
British Standards Institution. (2002). BS ISO 104:2002. Rolling bearings. Thrust bearings.
Boundary dimensions, general plan.
British Standards Institution. (2002). BS ISO 492:2002. Rolling bearings. Radial bearings.
Tolerances.
International Organization for Standardization. (1992). ISO 3448 Industrial liquid lubricants—
ISO viscosity classification.
International Organization for Standardization. (2001). ISO 14728-1.Rolling bearings. Linear
motion rolling bearings. Part 1: Dynamic load ratings and rating life.
International Organization for Standardization. (2001). ISO 14728-2.Rolling bearings. Linear
motion rolling bearings. Part 2: Static load ratings.
International Organization for Standardization. (2001). ISO 15243. Rolling bearings. Damages
and failures. Terms, characteristics and causes.
Society of Automotive Engineers. (1989). SAE J300. SAE standard engine oil viscosity
classification.
230
Mechanical Design Engineering Handbook
Further reading
Harnoy, A., 2002. Bearing Design in Machinery. Marcel Dekker, New York.
Hindhede, U., Zimmerman, J.R., Hopkins, R.B., Erisman, R.J., Hull, W.C., Lang, J.D., 1983.
Machine Design Fundamentals. A Practical Approach. Wiley, New York.
Neale, M.J. (Ed.), 1995. The Tribology Handbook. Butterworth Heinemann, Oxford.
Welsh, R.J., 1983. Plain Bearing Design Handbook. Butterworth, London, Boston.