IOP PUBLISHING JOURNAL OF MICROMECHANICS AND MICROENGINEERING
J. Micromech. Microeng. 17 (2007) 1257–1265 doi:10.1088/0960-1317/17/7/007
A micro electromagnetic generator for
vibration energy harvesting
S P Beeby
1
, R N Torah
1
, M J Tudor
1
, P Glynne-Jones
1
,
T O’Donnell
2
,CRSaha
2
and S Roy
2
1
School of Electronics and Computer Science, University of Southampton, Highfield,
Southampton, Hampshire, SO17 1BJ, UK
2
Tyndall National Institute, Prospect Row, Cork, Republic of Ireland
E-mail: spb@ecs.soton.ac.uk
Received 22 March 2007, in final form 6 May 2007
Published 5 June 2007
Online at stacks.iop.org/JMM/17/1257
Abstract
Vibration energy harvesting is receiving a considerable amount of interest as
a means for powering wireless sensor nodes. This paper presents a small
(component volume 0.1 cm
3
, practical volume 0.15 cm
3
) electromagnetic
generator utilizing discrete components and optimized for a low ambient
vibration level based upon real application data. The generator uses four
magnets arranged on an etched cantilever with a wound coil located within
the moving magnetic field. Magnet size and coil properties were optimized,
with the final device producing 46 µW in a resistive load of 4 k from just
0.59 m s
−2
acceleration levels at its resonant frequency of 52 Hz. A voltage
of 428 mVrms was obtained from the generator with a 2300 turn coil which
has proved sufficient for subsequent rectification and voltage step-up
circuitry. The generator delivers 30% of the power supplied from the
environment to useful electrical power in the load. This generator compares
very favourably with other demonstrated examples in the literature, both in
terms of normalized power density and efficiency.
(Some figures in this article are in colour only in the electronic version)
1. Introduction
Wireless sensor systems are receiving increasing interest
since they offer flexibility, ease of implementation and the
ability to retrofit systems without the cost and inconvenience
of cabling. Furthermore, by removing wires there is
the potential for embedding sensors in previously inaccessible
locations. At present, the majority of wireless sensor nodes
are simply battery-powered. Despite measures such as low
power techniques for communicating (e.g. IEEE 802.15.4
and Zigbee protocols) and the intelligent management of the
sensor node’s power consumption, batteries will still require
periodical replacement. Replacing batteries is not compatible
with embedded applications nor is it feasible for networks with
large numbers of nodes.
The advances made in low power wireless systems present
an opportunity for alternative types of power source. Solutions
such as micro fuel cells [1] and micro turbine generators
[2] are capable of high levels of energy and power density.
However, they involve the use of chemical energy and require
refuelling. Energy harvesting approaches that transform light,
heat and kinetic energy available in the sensor’s environment
into electrical energy offer the potential of renewable power
sources which can be used to directly replace or augment the
battery. Such renewable sources could increase the lifetime
and capability of the network and mitigate the environmental
impact caused by the disposal of batteries. In this context,
solar power is the most well known.
The subject of this paper is a kinetic energy generator
which converts mechanical energy in the form of vibrations
present in the application environment into electrical
energy. Kinetic energy is typically converted into electrical
energy using electromagnetic, piezoelectric or electrostatic
transduction mechanisms [3]. Vibrations are an attractive
source since the energy present can be harvested by compact
inertial devices that benefit from a high Q-factor amplifying the
base excitation amplitude. Suitable vibrations can be found in
numerous applications including common household goods
0960-1317/07/071257+09$30.00 © 2007 IOP Publishing Ltd Printed in the UK 1257
SPBeebyet al
Figure 1. Model of a linear, inertial generator.
(fridges, washing machines, microwave ovens), industrial
plant equipment, moving vehicles such as automobiles and
aeroplanes and structures such as buildings and bridges [4].
Human-based applications are characterized by low frequency
high amplitude displacements [5, 6].
The amount of energy generated by this approach
fundamentally depends upon the quantity and form of the
kinetic energy available in the application environment and
the efficiency of the generator and the power conversion
electronics. This paper presents the development of an
electromagnetic micro generator designed to harvest energy
from the vibrations of an air compressor unit which exhibits
large vibration maxima in the range of 0.19–3.7 m s
−2
at
frequencies between 43 Hz and 109 Hz. The micro generator
was therefore designed to operate within this range and to
be as small as possible whilst still generating useable levels
of power and voltage. The paper presents a theoretical
analysis of inertial generators, the design, simulation and
testing of the electromagnetic generator and a comparison
with other inertial generators. This work was carried out as
part of the European Union funded project ‘Vibration Energy
Scavenging’ (VIBES).
2. Basic theory
Resonant generators can be modelled as a second-order,
spring–mass–damper system with base excitation. Figure 1
shows a general example of such a system based on a seismic
mass, m, on a spring of stiffness, k. Total energy losses within
the system are represented by the damping coefficient, c
T
.
These losses consist of parasitic loss mechanisms (e.g. air
damping), represented by c
p
, and electrical energy extracted
by the transduction mechanism, represented by c
e
.
These generators are intended to operate at their resonant
frequency and for optimum energy extraction should be
designed such that this coincides with the vibrations present in
the intended application environment. The theory of inertial-
based generators is well documented [7–9] and will only be
briefly summarized here. Assuming the generator is driven
by a harmonic base excitation y(t) = Y sin(ωt), it will move
out of phase with the mass at resonance resulting in a net
displacement, z(t), between the mass and the frame.
The average power dissipated within the damper (i.e. the
power extracted by the transduction mechanism and the power
lost through parasitic damping mechanisms) is given by:
P
av
=
mξ
T
Y
2
ω
ω
n
3
ω
3
1 −
ω
ω
n
2
2
+
2ξ
T
ω
ω
n
2
(1)
where ξ
T
is the total damping ratio given by ξ
T
= c
T
/2mω
n
.
Since this equation is valid for steady-state conditions, P
av
is equal to the kinetic energy supplied per second by the
application vibrations. Maximum power dissipation within
the generator occurs when the device is operated at ω
n
and in
this case P
av
is given by:
P
av
=
mY
2
ω
3
n
4ξ
T
. (2)
Equation (2) suggests the following rules: (a) power varies
linearly with the mass; (b) power increases with the cube of
the frequency and (c) power increases with the square of the
base amplitude. Rules (b) and (c) are dependant upon the
base excitation, i.e. the accelerations present in the application
environment. Since the peak acceleration of the base, A,is
given by A = ω
2
Y and damping factor is related to the
damping ratio by c
T
= 2mω
n
ξ
T
, equation (2) can also be
written in the form
P
av
=
(
mA
)
2
2c
T
. (3)
These equations emphasize the need to understand the
vibrations present in the intended application when designing
an inertial generator. However, one cannot simply choose a
particular frequency of operation based upon the power output
alone. The inertial mass displacement will be limited to a given
finite value, z
max
, depending upon the size of the generator, its
design and material limitations. This is especially relevant in
the case of MEMS generators. Furthermore, z
max
will be a
multiple Q
T
times larger than Y where Q
T
is the total quality
factor of the generator given by equation (4):
Q
T
=
ω
n
m
c
T
=
1
2ξ
T
. (4)
The relationship between Q
T
and the electrical and parasitic
damping factors is given by equation (5)whereQ
OC
is the
open circuit Q-factor, i.e. 1/2ξ
P
,andQ
E
is equal to 1/2ξ
E
.
1
Q
T
=
1
Q
OC
+
1
Q
E
. (5)
Taking z
max
into consideration, average power can also be
expressed as
P
av
=
mω
3
n
Yz
max
2
. (6)
Incorporating the parasitic and electrical damping into
equation (2) gives the average power delivered to the electrical
domain:
P
avelec
=
mξ
E
Y
2
ω
3
n
4
(
ξ
P
+ ξ
E
)
2
. (7)
Maximum power is delivered to electrical domain when
ξ
E
= ξ
P
i.e. damping arising from the electrical domain should
equal mechanical losses. In this case equation (7) simplifies
to
P
avelec
=
mY
2
ω
3
n
16ξ
P
. (8)
Not all the energy transduced into the electrical domain
will actually be delivered into the load. In the case of
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A micro electromagnetic generator for vibration energy harvesting
0
0.01
0.02
0.03
0.04
0.05
0.06
0.07
0.08
0 20 40 60 80 100 120
Frequency, Hz
Acc, g
0.056g @ 43.3 Hz
0.035g @ 49.8 Hz
0.068g @ 108.8 Hz
0
0.05
0.1
0.15
0.2
0.25
0.3
0.35
0 25 50 75 100 125
Frequency, Hz
Acc, g
0.31g @ 49.7 Hz
Figure 2. Example vibration spectra from compressor application
(top plot from compressor enclosure, bottom plot from compressor).
electromagnetic transduction, some of the power delivered to
the electrical domain is lost within the coil. The actual power
in the load is a function of the coil and load resistances and is
calculated from equation (9).
P
L max
=
mω
3
n
Y
2
16ξ
P
R
load
R
load
+ R
coil
. (9)
However, the coil and load resistances also affect the damping
factor arising from electromagnetic transduction c
E
which can
be estimated from equation (10)whereN is the number of turns
in the generator coil, l is the side length of the coil (assumed
square) and B is the flux density to which it is subjected. R
L
,
R
coil
and L
coil
are the load resistance, coil resistance and coil
inductance respectively. Equation (12) is an approximation
and only ideal for the case where the coil moves in a region of
constant magnetic field.
c
E
=
(
NlB
)
2
R
L
+ R
coil
+ jωL
coil
. (10)
3. Application overview
The intended application for the generators described in this
paper is an air compressor unit supplying several laboratories
within a building. The electric motor runs continuously whilst
the compressor is duty cycled to maintain the pressure within
an in-line reservoir tank. The vibration levels and frequencies
have been measured at various locations on the compressor
and electric motor. The measured results indicate several
resonances between 43 and 109 Hz with acceleration levels
between 0.19 and 3.7 m s
−2
. Example vibration spectra taken
from the side of the compressor enclosure and the top of the
compressor are shown in figure 2. The generators presented
in this paper have been designed to operate at these lower
frequencies and at an rms acceleration of 0.59 m s
−2
(or
60 mg where 1 g = 9.81 m s
−2
). This frequency range and
Magnet
movement
Coil wire entering
Coil wire leaving page
Coil wire
Cantilever beam
Keeper
Magnets (poles shown)
N
N
S
S
S
S
N
N
Figure 3. Cross section through the four-magnet arrangement.
acceleration level is indicative of the vibration levels found in
typical industrial applications.
4. Mk1 electromagnetic generator design
4.1. Generator design overview
The micro electromagnetic generators presented in this paper
are a miniaturized form of a previous larger scale design
[10]. The generator uses miniature discrete components
fabricated using a variety of conventional manufacturing
processes. This enables the generator to exploit the advantages
of bulk magnetic material properties and large coil winding
density thereby demonstrating useable levels of power from a
compact design. A comparison between bulk and integrated
components for electromagnetic vibration energy harvesting
has been presented elsewhere [11].
The design uses four high energy density sintered rare
earth neodymium iron boron (NdFeB) magnets manually
bonded with Cyanoacrylate to the top and bottom surfaces
of a cantilever beam with the aid of an alignment jig. The
magnets were 1 × 1 × 1.5 mm
3
in size, being 1.5 mm in the
poled direction. The magnetic poles are aligned as shown in
figure 3. The magnetic circuit is completed by zinc coated
mild steel keepers which couple the flux between top and
bottom magnets. This arrangement produces a concentrated
flux gradient through the stationary coil as the magnets vibrate.
Additional mass is added to the generator in the form of
two wire eroded tungsten alloy blocks attached to the free end
of the cantilever beam. The tungsten alloy has a density of
18.1 g cm
−3
providing a compact inertial mass. The density
of the magnets is 7.6 g cm
−3
.
The beam used in this design was 9 mm long, 3 mm wide
along 7 mm of the beam length and 4 mm wide for the final
2 mm. Slots and holes have been incorporated into the beam to
accommodate the coil and bolt. All corners have radii to reduce
stress concentration effects. For the Mk1 generator, beams
were fabricated from double polished single crystal silicon
wafers. The geometry of the beam and required thickness was
determined by finite element analysis. A thickness of 50 µm
gave resonant frequencies between 50 and 60 Hz. Double
polished wafers were purchased in the desired thickness (with
a 5% tolerance), therefore having a high quality finish on both
top and bottom surfaces. The wafers were resist bonded to
a host wafer and the beams fabricated by deep reactive ion
etching through the 50 µm thickness.
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SPBeebyet al
Steel
washer
Tecatron GF40
base
Beam
Zintec keeper
Tungsten
mass
Copper
coil
NdFeB
magnets
Figure 4. Micro cantilever generator.
The cantilever beam assembly was clamped onto the base
using an M1 sized nut and bolt and a square washer. The
square washer gives a straight clamped edge perpendicular to
beam length. The base is machined from Tecatron GF40, a
40% glass fibre reinforced semi-crystalline high performance
plastic using a Daytron micro-mill. The high rigidity of the
material provides a firm clamping edge which is important
to avoid excessive energy loss through the fixed end of the
beam. The coil was manually bonded to a semi-circular recess
machined in the base. The coil has an outside radius R
o
of
1.2 mm, an inside radius R
i
of 0.3 mm and a thickness t of
0.5 mm. It was wound from 25 µm diameter enamelled
copper wire and had 600 turns. A drawing of the assembled
generator is shown in figure 4. With the aid of alignment
jigs, a tolerance of better than 0.1 mm can be achieved with
the manual assembly of the components. The volume of the
generator components is 0.1 cm
3
whilst the practical volume,
i.e. including the swept volume of the beam, is approximately
0.15 cm
3
.
4.2. Mk1 generator results
The generator produced a peak power of 10.8 µW from 60 mg
acceleration (1g = 9.81 m s
−2
) across a 110 load. The
voltage level generated was 34.5 mVrms. The generator
also demonstrated nonlinear behaviour which produced a
significant level of hysteresis in the output. This is shown
in figure 5 where the power output was measured as the
frequency was increased from below to above resonance and
also as the frequency was decreased from above to below
resonance. When reducing frequency the maximum power
that can be obtained is 2.5 µW. Whilst useable levels of power
were delivered to the load, the voltage level was too low to
enable subsequent voltage signal conditioning.
5. MK2 electromagnetic generator design
The generator was next subject to an optimization process with
the objectives of increasing the generated voltage and power
levels. In particular, the magnet size, beam material and coil
parameters were investigated
5.1. Finite element magnetic modelling
Ansoft Maxwell 3D magnetic finite element (FE) software was
used to optimize the electromagnetic circuit. The influence of
0
2
4
6
8
10
12
45 50 55 60
Frequency (Hz)
Power ( µW)
Increase Frequency
Decrease frequency
10.82µW@
58.5Hz
2.53µW
@54.9Hz
Figure 5. Power output hysteresis effect for the 50 µm beam micro
generator.
Figure 6. Magnet dimensions for simulation results.
magnet size was investigated by comparing the open circuit
voltage for various magnet widths and heights (dimensions x
and y respectively in figure 6). The thickness of the magnet,
w, was fixed at 1.5 mm and the distance between the magnets,
d, was fixed at 1 mm. The simulations were carried out with an
excitation frequency of 60 Hz, and acceleration of 0.59 m s
−2
.
Given a peak magnet amplitude of 0.57 mm, this corresponds
to a Q-factor of 140.
First, dimension y was fixed at 1 mm and x was varied
between 1 and 3 mm. The peak-generated voltage rises with
increasing x, but the rate of improvement reduces beyond
2.5 mm. Since, for a given volume, increasing magnet width
causes a reduction in the size of the proof mass, dimension
x was fixed at 2.5 mm. Next, with x fixed, y was adjusted
between 1 and 3 mm. The simulation results again show
an improvement in generated voltage with increasing y up to
2 mm. The simulation identified a practical optimum magnet
size of 2.5 × 2 × 1.5 mm
3
with further increases in magnet size
yielding diminishing improvements in voltage at the expense
of increased generator size and reduced mass. The predicted
voltage output for the increased magnet was 165 mVpk output
compared to 64 mVpk for the 1 × 1 × 1.5 mm
3
size magnets
(see figure 7). This is a factor of improvement of 2.6 in the
open circuit voltages.
5.2. Cantilever beam
Despite being an excellent spring material for this application,
the single-crystal silicon beams used in the Mk1 generator
were found to be too brittle to handle during assembly.
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A micro electromagnetic generator for vibration energy harvesting
Figure 7. Simulated output voltages for optimized and small
magnet generator configurations.
0
200
400
600
800
1000
1200
50 70 100 150 200 325
Beam Thickness ( m)
Frequency (Hz)
BeCu Si SS
Figure 8. Generator frequency for varying beam thickness and
material.
Therefore alternative metallic materials beryllium copper
(BeCu) and stainless steel type 302 full hard were investigated.
These materials possess mechanical properties well suited to
this application, in particular excellent fatigue characteristics.
The metal beams have been fabricated by a combination of
photolithography and spray etching. This involves coating
both sides of the metal sheet with a UV sensitive photoresist
and using contact lithography to define the beam shape.
After exposure, the resist is developed leaving regions of the
metal sheet exposed to a Ferric Chloride etchant which is
sprayed simultaneously to both sides. This etches through the
exposed metal leaving the desired beam geometry. This is a
straightforward batch fabrication process enabling numerous
structures to be fabricated simultaneously on each metal sheet.
The resonant frequency of the generator is defined by
the beam geometry, material and the inertial mass. The
resonant frequency of the generator versus beam thickness
is shown in figure 8 for a magnet size of 2.5 × 2 × 1.5 mm
3
.
These results were obtained from ANSYS modal analysis and
demonstrate the range of frequencies attainable with standard
sheet thicknesses. For this prototype 50 µm thick BeCu was
chosen which gives a predicted frequency of 51 Hz.
5.3. Coil properties
In addition to the coil used in the Mk1 generator two further
coils of identical dimensions were investigated. The three
coils, denoted by A, B and C, were wound from 25, 16 and
Table 1. Coil parameters.
Wire diameter, No. of Fill
Coil φ (µm) turns R
coil
() factor
A 25 600 100 0.67
B 16 1200 400 0.45
C 12 2300 1500 0.53
12 µm diameter enamelled copper wire respectively. Typical
coil parameters are given in table 1.
The length of wire used for each coil can be calculated
from L
w
= R
coil
A
w
/ρ where A
w
is the cross sectional area
of the wire and ρ the resistivity of copper (1.7 × 10
−8
m).
This gives wire lengths of 2.9, 4.7 and 10 m for coils A, B
and C respectively. The coil fill factor, F, the ratio of the
volume of conductor to the volume of the coil, is given by
equation (11):
F =
L
w
φ
2
4
R
2
o
− R
2
i
t
. (11)
This gives coil fill factors of 0.67 for coil A, 0.45 for coil B
and 0.53 for coil C. This shows there is a difference in the
density of the windings in each of the coils due to variations
in the winding process. A higher fill factor is preferable since
this indicates a higher number of turns within a given volume.
6. Experimental analysis
Testing of the generators was conducted using a shaker unit
with accelerometer feedback and a programmable resistive
load. The system is controlled by LabView software which
allows the user to program long sequences of tests to
automatically characterize each generator over a range of
acceleration levels, load resistances and frequencies. Great
care was taken to mount the accelerometer and generators
concentrically on the shaker unit to ensure reliable and
repeatable acceleration readings and results. The following
results were taken at an acceleration level of 60 mg, unless
otherwise noted.
6.1. Evaluation of optimized magnets
The first experiment compared the 1 × 1 × 1.5 mm
3
magnets
to the optimized dimensions, 2.5 × 2 × 1.5 mm
3
using coil A.
The comparison is shown in figure 9 which shows the measured
voltage across a 9 M load resistance versus frequency.
The observed resonant frequency of 56.6 Hz shows
reasonable agreement with the FEA model being within 10%
of the predicted result. The difference is due to the tolerance
on the thickness of the beam and the nonlinear response of the
generator. The peak output voltage increases from 39 mVrms
with the original magnets to 88 mVrms with the optimized
magnet configuration, an increase of 225%. Next, the power
output to the load was measured for the optimized magnet
configuration. The optimum load resistance was determined
by measuring the power output at resonance over a wide range
of resistance values, the optimum being 150 . The maximum
power output of 17.8 µW was obtained with a voltage output
of 52 mVrms across the optimum load as shown in figure 10.
Both figures 9 and 10 show evidence of the magnets on the
beam touching the base at peak amplitude. The base was
modified to avoid this in subsequent experiments.
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