A micro electromagnetic generator for vibration energy harvesting

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A micro electromagnetic generator for vibration energy harvesting
This article has been downloaded from IOPscience. Please scroll down to see the full text article. 2007 J. Micromech. Microeng. 17 1257 (https://www.wendangku.net/doc/8510442404.html,/0960-1317/17/7/007) View the table of contents for this issue, or go to the journal homepage for more
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IOP PUBLISHING J. Micromech. Microeng. 17 (2007) 1257–1265
JOURNAL OF MICROMECHANICS AND MICROENGINEERING
doi:10.1088/0960-1317/17/7/007
A micro electromagnetic generator for vibration energy harvesting
S P Beeby1, R N Torah1, M J Tudor1, P Glynne-Jones1, T O’Donnell2, C R Saha2 and S Roy2
1
School of Electronics and Computer Science, University of Southampton, High?eld, Southampton, Hampshire, SO17 1BJ, UK 2 Tyndall National Institute, Prospect Row, Cork, Republic of Ireland E-mail: spb@https://www.wendangku.net/doc/8510442404.html,
Received 22 March 2007, in ?nal form 6 May 2007 Published 5 June 2007 Online at https://www.wendangku.net/doc/8510442404.html,/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 cm3, practical volume 0.15 cm3) 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 ?eld. Magnet size and coil properties were optimized, with the ?nal 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 suf?cient for subsequent recti?cation 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 ef?ciency. (Some ?gures in this article are in colour only in the electronic version)
1. Introduction
Wireless sensor systems are receiving increasing interest since they offer ?exibility, ease of implementation and the ability to retro?t 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.
0960-1317/07/071257+09$30.00 ? 2007 IOP Publishing Ltd
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 bene?t from a high Q-factor amplifying the base excitation amplitude. Suitable vibrations can be found in numerous applications including common household goods 1257
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S P Beeby et al
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: Pav = mξT Y 2 1?
ω 2 2 ωn ω 3 3 ω ωn ω ωn 2
(1)
+ 2ξT
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 ef?ciency 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 coef?cient, cT. These losses consist of parasitic loss mechanisms (e.g. air damping), represented by cp , and electrical energy extracted by the transduction mechanism, represented by ce . 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 inertialbased generators is well documented [7–9] and will only be brie?y 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. 1258
where ξT is the total damping ratio given by ξT = cT /2mωn . Since this equation is valid for steady-state conditions, Pav 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 ωn and in this case Pav is given by: 3 mY 2 ωn Pav = . (2) 4ξT 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 = ω2 Y and damping factor is related to the damping ratio by cT = 2mωn ξT , equation (2) can also be written in the form (mA)2 . (3) Pav = 2cT 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 ?nite value, zmax , depending upon the size of the generator, its design and material limitations. This is especially relevant in the case of MEMS generators. Furthermore, zmax will be a multiple QT times larger than Y where QT is the total quality factor of the generator given by equation (4): ωn m 1 QT = = . (4) cT 2ξT The relationship between QT and the electrical and parasitic damping factors is given by equation (5) where QOC is the open circuit Q-factor, i.e. 1/2ξ P, and QE is equal to 1/2ξ E. 1 1 1 = + . (5) QT QOC QE Taking zmax into consideration, average power can also be expressed as 3 mωn Y zmax . (6) Pav = 2 Incorporating the parasitic and electrical damping into equation (2) gives the average power delivered to the electrical domain: 3 mξE Y 2 ωn . (7) Pavelec = 4 (ξP + ξE )2 Maximum power is delivered to electrical domain when ξE = ξP i.e. damping arising from the electrical domain should equal mechanical losses. In this case equation (7) simpli?es to 3 mY 2 ωn . (8) Pavelec = 16ξP Not all the energy transduced into the electrical domain will actually be delivered into the load. In the case of

A micro electromagnetic generator for vibration energy harvesting
0.08 0.07 0.06
0.056g @ 43.3 Hz 0.068g @ 108.8 Hz
S
0.035g @ 49.8 Hz
N
S
N
Acc, g
0.05 0.04 0.03 0.02 0.01 0 0 20 40 60 80 100 120
Magnet movement
N
Coil wire entering
S
N
S
Cantilever beam Keeper Magnets (poles shown)
Frequency, Hz
0.35 0.3 0.25
Coil wire leaving page Coil wire
0.31g @ 49.7 Hz
Figure 3. Cross section through the four-magnet arrangement.
Acc, g
0.2 0.15 0.1 0.05 0 0 25 50 75 100 125
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 mm3 in size, being 1.5 mm in the poled direction. The magnetic poles are aligned as shown in ?gure 3. The magnetic circuit is completed by zinc coated mild steel keepers which couple the ?ux between top and bottom magnets. This arrangement produces a concentrated ?ux 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 ?nal 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 ?nite 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 ?nish 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. 1259
Frequency, 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).
3 Rload mωn Y 2 . (9) 16ξP Rload + Rcoil However, the coil and load resistances also affect the damping factor arising from electromagnetic transduction cE which can be estimated from equation (10) where N is the number of turns in the generator coil, l is the side length of the coil (assumed square) and B is the ?ux density to which it is subjected. RL, Rcoil and Lcoil 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 ?eld.
PL max =
cE =
(N lB)2 . RL + Rcoil + j ωLcoil
(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 ?gure 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

S P Beeby et al
NdFeB magnets
Power (μW ) 12 10 8 6 4 2 0 45 50 55 Frequency (Hz) 60 2.53μW @54.9Hz Increase Frequency Decrease frequency 10.82μW @ 58.5Hz
Steel washer
Copper coil
Tecatron GF40 base
Beam
Zintec keeper
Tungsten mass
Figure 5. Power output hysteresis effect for the 50 μm beam micro generator.
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 ?bre reinforced semi-crystalline high performance plastic using a Daytron micro-mill. The high rigidity of the material provides a ?rm clamping edge which is important to avoid excessive energy loss through the ?xed end of the beam. The coil was manually bonded to a semi-circular recess machined in the base. The coil has an outside radius Ro of 1.2 mm, an inside radius Ri 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 ?gure 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 cm3 whilst the practical volume, i.e. including the swept volume of the beam, is approximately 0.15 cm3. 4.2. Mk1 generator results The generator produced a peak power of 10.8 μW from 60 mg load. The acceleration (1g = 9.81 m s?2) across a 110 voltage level generated was 34.5 mVrms. The generator also demonstrated nonlinear behaviour which produced a signi?cant level of hysteresis in the output. This is shown in ?gure 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.
Figure 6. Magnet dimensions for simulation results.
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 ?nite element (FE) software was used to optimize the electromagnetic circuit. The in?uence of 1260
magnet size was investigated by comparing the open circuit voltage for various magnet widths and heights (dimensions x and y respectively in ?gure 6). The thickness of the magnet, w, was ?xed at 1.5 mm and the distance between the magnets, d, was ?xed 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 ?xed 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 ?xed at 2.5 mm. Next, with x ?xed, 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 identi?ed a practical optimum magnet size of 2.5 × 2 × 1.5 mm3 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 mm3 size magnets (see ?gure 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.

A micro electromagnetic generator for vibration energy harvesting
Table 1. Coil parameters. Coil A B C Wire diameter, φ (μm) 25 16 12 No. of turns 600 1200 2300 Rcoil ( ) 100 400 1500 Fill factor 0.67 0.45 0.53
Figure 7. Simulated output voltages for optimized and small magnet generator con?gurations.
1200 1000
Frequency (Hz)
BeCu 800 600 400 200 0 50
Si
SS
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 Lw = Rcoil Aw /ρ where Aw 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 ?ll factor, F, the ratio of the volume of conductor to the volume of the coil, is given by equation (11): Lw φ 2 . (11) F = 2 4 Ro ? Ri2 t This gives coil ?ll 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 ?ll factor is preferable since this indicates a higher number of turns within a given volume.
6. Experimental analysis
70 100 150 200 325
Beam Thickness ( m)
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 de?ne 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 de?ned by the beam geometry, material and the inertial mass. The resonant frequency of the generator versus beam thickness is shown in ?gure 8 for a magnet size of 2.5 × 2 × 1.5 mm3. 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
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 ?rst experiment compared the 1 × 1 × 1.5 mm3 magnets to the optimized dimensions, 2.5 × 2 × 1.5 mm3 using coil A. The comparison is shown in ?gure 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 con?guration, an increase of 225%. Next, the power output to the load was measured for the optimized magnet con?guration. 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 ?gure 10. Both ?gures 9 and 10 show evidence of the magnets on the beam touching the base at peak amplitude. The base was modi?ed to avoid this in subsequent experiments. 1261

S P Beeby et al
100
500 600turn
Load Voltage (mVrms)
80 60 40 20 0 53 54 55 56 57 58
Original magnets
Voltage (mVrms)
400
1200turn 2300turn
Optimized magnets
300
200
100
59
60
61
62
63
0 48 49 50 51 52 53 54 55
Frequency (Hz)
Frequency (Hz)
Figure 9. Generated output voltages for the small and optimized magnet con?gurations.
20
Figure 11. Output voltages for 600/1200/2300 turn coils across optimum load resistance.
50 600turn
15
Power ( Wrms)
40
1200turn 2300turn
Power (uW)
30
10
20
5
10
0 46 51 56 61
0 48 49 50 51 52 53 54 55
Frequency (Hz)
Frequency (Hz)
Figure 10. Optimum power output of the optimized magnet con?guration generator. Table 2. Generator results. Coil A B C Resonant frequency 52.1 Hz 51.64 Hz 53.2 Hz Load resistance 200 500 4k Power at resonance 44.7 μWrms 45.8 μWrms 45.7 μWrms Voltage at resonance 95 mVrms 151 mVrms 428 mVrms
Figure 12. Output power for 600/1200/2300 turn coils with optimum load resistance.
6.2. Evaluation of coil types Coils of types A, B and C were each located on an individual generator base and the same beam assembly mounted on each base in turn. The resulting output voltage across the optimum load for each coil type is shown in ?gure 11. As expected, the output voltage increases with increasing number of turns with 95, 151 and 428 mVrms being generated from the 600, 1200 and 2300 turn coils respectively. The generated power is very similar for each device as shown in ?gure 12 with the full set of results summarized in table 2. The generated power is essentially the same for each device because the improved voltage output is offset by the increased coil resistance. This is re?ected by the equation for damping factor (equation (10)). The increase in voltage for the generator with coil C to over 400 mVrms should be suf?cient to enable conventional passive recti?cation and step up circuits to be implemented. Furthermore, the generator output power has increased to over 45 μW at 60 mg excitation. This is due to improvements in the assembly of the device, in particular the clamping and alignment of the beam, which leads to reduced energy parasitic damping and an increased open circuit Q-factor. 1262
The high QO/C means these generators demonstrated nonlinear behaviour at very low acceleration levels (<3 mg) and the Q-factor cannot be determined from a frequency amplitude plot such as that shown in ?gure 12. Therefore, the Q-factor was measured by observing the decay in the voltage signal from the generator. The coil C generator was initially driven at resonance at 20 mg and the excitation stopped abruptly. Figure 13(a) shows the decay in the generator open circuit output over 6.5 s, ?gure 13(b) shows the decay immediately after turning off the excitation and ?gure 13(c) shows the signal approximately 1.5 s later. Q-factor can be calculated from equation (12) where f0 is the frequency, V1 and V2 are the voltage amplitudes at a time interval t apart: Q= πf0 t ln
V1 V2
.
(12)
The Q-factor calculated from ?gure 13(b) is 520 whilst the Q-factor in (c) is 274. This indicates that the parasitic damping is a function of amplitude and decreases with increasing excitation acceleration levels. This behaviour could be due to the magnets extending beyond the in?uence of the coil at the higher amplitudes which, even when open circuit, produces a damping effect.
7. Theoretical analysis of device performance
Given the nonlinear behaviour of the high-Q generators, the generator used to evaluate the optimum magnet size (described

A micro electromagnetic generator for vibration energy harvesting
(a) 0.8
Power amp switched off
0.6 0.4
14 12
Power (uW)
Measured power across load Input Power (equ. 2) Power in electrical domain (equ. 8) Power in the load (equ. 9)
10 8 6 4 2
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0 1 2 3 4 5 6
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5
10
15
20
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30
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Figure 14. Comparison of theoretical and measured power output.
Voltage (V)
0.2 0 0.8 -0.2 -0.4 -0.6 -0.8 0.85 0.9 0.95 1 1.05 1.1 1.15
Time (s)
(c ) 0.8
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have been plotted versus acceleration up to 0.29 m s?2 in ?gure 14. This shows excellent agreement between measured and predicted power levels and demonstrates that the cantilever microgenerator is converting 30% of the total power dissipated in the generator to electrical power delivered to the load. Due to the relative coil and load resistances, one third of the power converted into the electrical domain is lost within the coil. Equation (9) highlights the importance of reducing coil resistance and increasing load resistance as long as the optimum damping condition is maintained. The theoretical and practical results show excellent agreement up to 0.29 m s?2. Beyond this point, the nonlinear behaviour alters the frequency response and parasitic damping levels and determining accurate values for the theoretical analysis is not straightforward. Work to analyse the theoretical response of the high-Q generators in the nonlinear region is ongoing.
8. A comparison of selected energy harvesting devices
Time (s)
Figure 13. Decay plots from generator with coil C (a) over 6.5 s (top), (b) initial decay after switching shaker off, Q = 520 (middle), (c) decay 1.5 s later, Q = 274 (bottom).
in section 6.1, peak power 17.8 μW at 56.6 Hz) was compared with the theory. This device was more highly damped and did not demonstrate nonlinear frequency or nonlinear damping behaviour below 0.29 m s?2. The generator was compared with the theoretical calculations for power supplied to the generator from the environment (equation (2)), the maximum power delivered to the electrical domain (equation (8)) and the power delivered to the load (equation (9)). Firstly, the closed and open loop quality factors were measured at an acceleration level below the onset of nonlinear behaviour. The measured values of QT and QOC at 20 mg were 119 and 232 respectively and from equation (5), QE equals 243. It can be seen that the optimum damping conditions have very nearly been met and it is reasonable to use equations (8) and (9). The coil and load resistances, RCoil and RLoad, used in equation (9) were equal to 100 and 200 respectively. The predicted power outputs from the theoretical equations and the measured output from the generator
Comparing different vibration energy harvesters is not straightforward since the amount of data presented in published works varies considerably. Therefore, inevitably some factors have to be extrapolated from the data given and any comparison should only be treated as a guide. Mitcheson et al [12] presented a comparison where they estimated the relative input powers and calculated an approximate ef?ciency. In this paper, we have derived a ?gure for normalized power density (NPD) which is simply the stated power output of the device normalized to acceleration level and divided by the volume. Frequency is not considered since resonant generators are ?xed in frequency whereas acceleration levels applied during testing can be varied. Neither the ef?ciency nor NPD metric is ideal since they both ignore important factors such as bandwidth but unfortunately insuf?cient data exist in the literature to enable this to be included. The Perpetuum generator [14], for example, has a signi?cantly broader bandwidth than the generator presented here enabling it to harvest energy from a wider range of frequencies. Nonetheless, the comparison of NPD for different devices can provide an indication of relative performance levels and provide a useful insight into trends and advances. Since power output varies with acceleration2, the calculated NPD is given by P /A2 V where P is the stated power 1263

S P Beeby et al
1000.00
VIBES Mk 2 (EM)
100.00
Glynne-Jones (EM) Perpetuum PMG7 (EM)
NPD (kgs/m3)
10.00 1.00 0.10 0.01 0.00 0.00001 0.0001
Ching (EM) Glynne-Jones (PZ) Roundy (PZ) Hong (PZ) Jeon (PZ) Mitcheson (ES) Despesse (ES)
0.001
0.01
0.1
1
10
100
Volume (cm3)
Figure 15. Normalized power density versus device volume. Table 3. Comparison of generators. Generator
a
Freq (Hz) 52 99 100 110 80 120 190 13 900 30 50
Acceln (m s?2) 0.589 6.85 0.400 95.5 2.3 2.5 71.3 106.8 50 8.8
Inertial mass (g) 0.66 2.96 50 0.192 0.8 9.15 0.01 2.20 × 10?07 0.1 104
Volume (cm3) 0.15 4.08 30 1 0.125 1 0.0012 0.000 027 0.75 1.8
Power (μW) 46 4990 4000 830 2.1 375 65 1 3.7 1052
NPD (kgs m?3) 883.97 26.07 833.33 0.09 3.18 60.00 10.67 3.25 0.002 7.55
VIBES Mk2 EM Glynne-Jones [13] EM Perpetuum [14] EM Ching [15] EM White [16] PZ Roundy [17] PZ Hong [18] PZ Jeon [19] PZ Mitcheson [20] ES Despesse [21] ES
a
Generators are labelled by technology: EM, electromagnetic; PZ, piezoelectric; ES, electrostatic.
output and V is the reported volume of the generator. The selection and details of energy harvesters from the literature used in the comparison are detailed in table 3 and a comparison is shown diagrammatically in ?gure 15.
9. Conclusions
This paper has presented a small (components volume 0.1 cm3, practical volume 0.15 cm3) electromagnetic vibration energy harvesting device optimized for a low level of ambient vibration based upon real application data. It is capable of producing useful power from a vibration level of 60 mg, delivering 46 μW to a resistive load of 4 k when the device is shaken at its resonant frequency of 52 Hz. This is a power density of 307 μW m?3. The generator delivers 30% of the total power dissipated in the generator to 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 ef?ciency. From the basic equations governing electromagnetic generators it is clear that the generated energy decreases with device volume, and reducing input vibration acceleration2. With the exception of the Perpetuum generator [14] (which is a much larger generator, 30 cm3), all the existing generators (both piezoelectric and electromagnetic) described in the literature [3] produce optimum power densities at input vibrations greater than 2 m s?2. One of the aims of this study was to demonstrate that it was possible for a device of this volume to produce useful power from only 0.59 m s?2 of ambient vibration. 1264
To demonstrate that the power from the device is ‘useful’ (i.e. of a voltage level and source impedance that permits recti?cation and storage), a prototype recti?cation and multiplier circuit has been ?tted to the generator (in a manner similar to that described by Ching et al [15]), and is capable of charging a capacitor to power the transmission of a periodic radio signal. This circuitry was not optimized, and we continue to work on a demonstrator that will better show the potential of a miniature wireless sensor node based on this generator.
Acknowledgments
The authors would like to acknowledge the ?nancial support of the European Union who funded this research through the Framework 6 Programme. We would like to acknowledge Dr Steve Roberts at Perpetuum Ltd for his assistance in setting up the vibration test rig.
References
[1] Banazwski B and Shah R K 2003 The role of fuel cells for consumer electronic products and toys Proc. 1st Int. Conf. on Fuel Cell Science, Engineering and Technology (Rochester, NY) pp 149–55 [2] Epstein A H 2004 Millimeter-scale, micro-electro-mechanical systems gas turbine engines J. Eng. Gas Turbines Power 126 205–26 [3] Beeby S P, Tudor M J and White N M 2006 Energy harvesting vibration sources for microsystems applications Meas. Sci. Technol. 17 R175–R195

A micro electromagnetic generator for vibration energy harvesting
[4] Roundy S, Wright P K and Rabaye J 2003 A study of low level vibrations as a power source for wireless sensor nodes Comput. Commun. 26 1131–44 [5] Starner T and Paradiso J A 2004 Human generated power for mobile electronics Low Power Electronics Design ed C Piguet (Boca Raton, FL: CRC Press) [6] von B¨ ren T, Lukowicz P and Tr¨ ster G 2003 Kinetic energy u o powered computing—an experimental feasibility study Proc. 7th IEEE Int. Symp. on Wearable Computers ISWC ’03 (White Plains, NY) pp 22–4 [7] El-Hami M et al 2001 Design and fabrication of a new vibration-based electromechanical power generator Sensors Actuators A 92 335–42 [8] Williams C B, Shearwood Harradine M A, Mellor P H, Birch T S and Yates R B 2001 Development of an electromagnetic micro-generator IEE Proc. Circuits Devices Syst.148 337–42 [9] Stephen N G 2006 On energy harvesting from ambient vibration J. Sound Vib. 293 409–25 [10] Glynne-Jones P, Tudor M J, Beeby S P and White N M 2004 An electromagnetic, vibration-powered generator for intelligent sensor systems Sensors Actuators A 110 344–49 [11] O’Donnell T, Saha C, Beeby S and Tudor J 2006 Scaling effects for electromagnetic vibrational power generators Proc. Design, Test, Integration and Packaging of MEMS and MOEMS (Stresa, Italy, 26–28 April 2006) Proc. SPIE 4019 279–85 [12] Mitcheson P D, Reilly E K, Wright P K and Yeatman E M 2006 Transduction mechanisms and power density for MEMS inertial energy scavengers Proc. Power MEMS 06 (Berkeley: USA) pp 275–78
[13] Glynne-Jones P 2001 Vibration powered generators for self-powered microsystems PhD Dissertation Department of Electronics and Computer Science, University of Southampton, UK [14] PMG7 data sheet, Perpetuum Ltd. https://www.wendangku.net/doc/8510442404.html,. [15] Ching N N H, Wong H Y, Li W J, Leong P H W and Wen Z 2002 A laser-micromachined vibrational to electrical power transducer for wireless sensing systems Sensors Actuators A 97–98 685–90 [16] Glynne-Jones P, Beeby S P and White N M 2001 Towards a piezoelectric vibration powered microgenerator IEE Proc. Sci. Meas. Technol. 148 68–72 [17] Roundy S and Wright P K 2004 A piezoelectric vibration based generator for wireless electronics Smart Mater. Struct. 13 1131–42 [18] Hong Y K and Moon K S 2005 Single crystal piezoelectric transducers to harvest vibration energy Proc. SPIE 6048 60480E-1 [19] Jeon Y B, Sood R, Jeong J-h and Kim S G 2005 MEMS power generator with transverse mode thin ?lm PZT Sensors Actuators A 122 16–22 [20] Mitcheson P, Stark B, Miao P, Yeatman E, Holmes A and Green T 2003 Analysis and optimisation of MEMS on-chip power supply for self powering of slow moving sensors Proc. Eurosensors XVII (Guimaraes: Portugal) pp 30–1 [21] Despesse G et al 2005 Fabrication and characterisation of high damping electrostatic micro devices for vibration energy scavenging Proc. Design, Test, Integration and Packaging of MEMS and MOEMS pp 386–90
1265

常用贴片元件封装尺寸图

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封装形式图片国际统一简称 LDCC LGA LQFP PDIP TO5 TO52 TO71 TO71 TO78 PGA Plastic PIN Grid Array 封装形式图片国际统一简称 TSOP Thin Small OUtline Package QFP Quad Flat Package PQFP 100L QFP Quad Flat Package SOT143 SOT220 Thin Shrink Qutline Package uBGA Micro Ball Grid Array uBGA Micro Ball Grid Array PCDIP

PLCC LQFP LQFP 100L TO8 TO92 TO93 T099 EBGA 680L QFP Quad Flat Package TQFP 100L ZIP Zig-Zag Inline Packa SOT223 SOT223 SOT23 SOT23/SOT323 SOT25/SOT353 SOT26/SOT363 FBGA FDIP SOJ

SBGA LBGA 160L PBGA 217L Plastic Ball Grid Array SBGA 192L TSBGA 680L CLCC SC-705L SDIP SIP Single Inline Package SO Small Outline Package SOP EIAJ TYPE II 14L SSOP 16L SSOP SOJ 32L Flat Pack HSOP28 ITO220 ITO3P TO220 TO247

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接触器文字符号位：KM 接触器图形符号位： 3 热继电器 热继电器时利用电流通过元件产生的热效应原理而反时限动作的继电器。

热继电器的文字符号位：FR 热继电器的图形符号位： 4 中间继电器 中间继电器的原理是将一个输入信号变成多个输出信号或将信号放大（即增大继电器触头容量）的继电器。其实质是电压继电器，但他的触头较多（可多达8对）、触头容量可达5-10A、动作灵敏。 当其他电器的触头对数不够时，可借助中间继电器来扩展他们的触头对数，也有通过中间继电器实现触电通电容量的扩展。

中间继电器的文字符号位：KA 中间继电器的图形符号位：

5 按钮 在实际应用中通常根据所需的触头数量、使用的场合及颜色来选择按钮。通常用的LA18、LA19、LA20等系列按钮，适用于AC500V、DC440V、额定电流5A，控制功率在AC300W、DC70W的控制回路中。 按钮文字符号位：SB 按钮图形符号位：

按钮颜色要求： 1）“停止”和“急停”按钮必须是红色。当按下红色按钮时必须使用设备停止运行或断电。2）“启动按钮”的颜色是绿色。 3）“启动”和“停止”交替动作的按钮必须是黑色、白色或灰色，不得使用红色和绿色按钮。 4）“点”动的按钮必须是黑色。 5）“复位”（如果有保护继电器的复位按钮）必须是蓝色，当复位按钮同时还有停止作用时，则必须是红色。 6 指示灯 指示灯的作用： 1）指示设备的运行或停止状态。 2）监视控制电器的电源是否正常， 3）利用红灯监控跳闸回路是否正常，用绿灯监控合闸回路是否正常。 7 旋转开关 万能转化开关由操作机构、面板、手柄、和数个触头座等部件组成。

常用贴片元件封装尺寸图

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(整理)常用PCB封装图解

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最新SMT常见贴片元器件封装类型和尺寸资料

1、SMT表面封装元器件图示索引（完善）

2、SMT物料基础知识 一. 常用电阻、电容换算： 1.电阻（R）： 电阻：定义：导体对电流的阻碍作用就叫导体的电阻。 无方向,用字母R表示,单位是欧姆(Ω)，分：欧（Ω）、千欧（KΩ）、兆欧（MΩ）1MΩ=1000KΩ=1000000Ω 1).换算方法： ①.前面两位为有效数字(照写)，第三位表示倍数10n次方（即“0”的个数） 103=10*103=10000Ω=10KΩ 471=47*101=470Ω 100=10*100=10Ω 101=10×101=100Ω 120=12×100=12Ω ②.前面三位为有效数字(照写)，第四位表示倍数倍数10n次方（即“0”的个数）. 1001=100*101=1000Ω=1KΩ 1632=163*102=16300Ω=16.3KΩ 1470=147×100=147Ω 1203=120×103Ω=120KΩ 4702=470×102Ω=47KΩ

2.电容（C）： 电容的特性是可以隔直流电压，而通过交流电压。它分为极性和非极性，用C表示。 2.1三种类型：电解电容钽质电容有极性, 贴片电容无极性。 用字母C表示，单位是法（F），毫法（MF），微法（UF）,纳法（NF）皮法（PF） 1F=103MF=106UF=109NF=1012PF 2.2换算方法: 前面两位为有效数字(照写)，第三位倍数10n次方（即“0”的个数） 104=10*104=100000PF=0.1UF 100=10*100=10PF 473=47×103=47000pF=47nF=0.047uF 103=10×103=10000pF=10nF=0.01uF 104=10×104=100000pF=10nF=0.1uF 221=22×101=220pF 330=33×100=33pF 2.3钽电容： 它用金属钽或者铌做正极，用稀流酸等配液做负极，用钽或铌表面生成的氧化膜做成介质制成，其特点是体积小、容量大、性能稳定、寿命长、绝缘电阻大、温度特性好，用在要求较高的设备中。钽电容表面有字迹表明其方向、容值，通常有一条横线的那边标志钽电容的正极。钽电容规格通常有：A型、B型、C型、P型。 2.4 电容的误差表示 2.4.1常用钽电容代换参照表. 1UF：105、A6、CA6 2.2UF：225 3.3UF：335、AN6、CN6、JN6、CN69 4.7UF：475、JS6 10UF：106、JA7、AA7、GA7 22UF：226、GJ7、AJ7、JJ7 47UF：476 3. 电感（L）

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常见元件的封装形式.

常见元件的封装形式 对于集成电路芯片来说，常见的封装形式有DIP(即双列直插式)，根据封装材料的不同，DIP封装又可以分为PDIP(塑料双列直插式)和CDIP(陶瓷双列直插式)； SIP(即单列直插式)；SOP(即小尺寸封装)；PQFP(塑料四边引脚扁乎封装)；PLCC(塑料有引线芯片载体封装)和LCCC(陶瓷无引线芯片载体封装)；PGA(插针网格阵列)、BGA(球形网格阵列)等。对于具体型号的集成电路芯片来说，其封装形式是固定的，如74系列集成电路芯片，一般采用DIP 封装形式，只有个别芯片生产厂家提供两种或两种以上封装形式。 对于分立元件(如电阻、电容、电感)来说，元件封装尺寸与元件大小、耗散功率、安装方式等因素有关。 1.电阻器常用的封装形式 电阻器常用的封装形式是AXIALO·3～AXIALl·0，对于常用的1／81r小功率电阻来说，可采用AXlALO·3或AXIALO·4(即两引线孔间距为0·762～1·016cm)：对于1／4W电阻来说，可采用AXlALO·5(即两引线孔间距为1·27cm)，但当采用竖直安装时，可采用AXlALO·3。 2.小容量电解电容的封装形式 小容量电解电容的封装形式一般采用RB·2／．4(两引线孔距离为0·-2英寸，而外径为0．4英寸)到RB．5／1·0(两引线孔距离为O·5英寸，而外径为1·0英寸)。对于大容量电容，其封装尺寸应根据实际尺寸来决定。 3.普通二极管封装形式 普通二极管封装形式为DIODEO·4～DIODEO·7。 4.三极管的封装形式 三极管的封装形式由三极管型号决定，常见的有T0-39、T0-42、T0-54、TO-92A、TO-92B、T0-220等。 对于小尺寸设备，则多采用表面安装器件，如电阻、电容、电感等一股采用SMC线封装方式；对于三极管、集成电路来说，多采用SMD封装方式。 进行电原理图编辑和印制板设计时，器件型号完全确定后，可以从器件手册查到封装形式和尺寸，或测绘后用PCBLib编辑器创建元件的封装图。 当用户实在无法确定时，也可以在PCB编辑器窗口内，单击元件“放置工具"，再 单击“测览”按钮，从Advpcb·ddb元件封装图形库中找出所需元件的封装形式。

常用电子元件知识分类及工作原理

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贴片元器件封装尺寸

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常用贴片元件封装

常用贴片元件封装 1 电阻： 最为常见的有0201、0402、0805、0603、1206、1210、1812、2010、2512几类1）贴片电阻的封装与尺寸如下表： 英制(mil) 公制(mm) 长(L)(mm) 宽(W)(mm) 高(t)(mm) 0201 0603 ±±± 0402 1005 ±±± 0603 1608 ±±± 0805 2012 ±±± 1206 3216 ±±± 1210 3225 ±±± 1812 4832 ±±± 2010 5025 ±±± 2512 6432 ±±± 2）贴片电阻的封装、功率与电压关系如下表： 英制(mil)公制(mm)额定功率@ 70°C 最大工作电压(V) 0201 0603 1/20W 25 0402 1005 1/16W 50 0603 1608 1/10W 50 0805 2012 1/8W 150 1206 3216 1/4W 200 1210 3225 1/3W 200 1812 4832 1/2W 200 2010 5025 3/4W 200 2512 6432 1W 200 3）贴片电阻的精度与阻值 贴片电阻阻值误差精度有±1％、±2％、±5％、±10％精度， J －表示精度为5％、 F－表示精度为1％。 T －表示编带包装 阻值范围从0R-100M 4）贴片电阻的特性 ·体积小，重量轻； ·适应再流焊与波峰焊； ·电性能稳定，可靠性高； ·装配成本低，并与自动装贴设备匹配； ·机械强度高、高频特性优越。 2电容： 1）贴片电容可分为无极性和有极性两种，容值范围从 无极性电容下述两类封装最为常见，即0805、0603； 英制尺寸公制尺寸长度宽度厚度 0402 1005 ± ± ± 0603 1608 ± ±±

常用贴片元件封装.

常用贴片元件封装 1 电阻： 最为常见的有 0201 、040 2 、 0805 、060 3 、1206 、 1210 、1812 、2010 、2512 几类 1)贴片电阻的封装与尺寸如下表： 英制(mil)公制(mm)长(L)(mm)宽(W)(mm) 高(t)(mm) 英制(mil)公制(mm)额定功率@ 70 C 最大工作电压(V) 0201 0603 1/20W 25 0402 1005 1/16W 50 0603 1608 1/10W 50 0805 2012 1/8W 150 **** **** 1/4W 200 1210 3225 1/3W 200 1812 4832 1/2W 200 0201 0603 0.60 ±0.05 0.30 ±0.05 0.23 ±0.05 0402 1005 1.00 ±0.10 0.50 ±0.10 0.30 ±0.10 0603 1608 1.60 ±0.15 0.80 ±0.15 0.40 ±0.10 0805 2012 2.00 ±0.20 1.25 ±0.15 0.50 ±0.10 1206 3216 3.20 ±0.20 1.60 ±0.15 0.55 ±0.10 1210 3225 3.20 ±0.20 2.50 ±0.20 0.55 ±0.10 1812 4832 4.50 ±0.20 3.20 ±0.20 0.55 ±0.10 2010 5025 5.00 ±0.20 2.50 ±0.20 0.55 ±0.10 2512 6432 6.40 ±0.20 3.20 ±0.20 0.55 ±0.10 功率与电压关系如下表： 2)贴片电阻的封装、

贴片元件常见封装

贴片电阻常见封装有9种，用两种尺寸代码来表示。一种尺寸代码是由4位数字表示的EIA(美国电子工业协会)代码，前两位与后两位分别表示电阻的长与宽，以英寸为单位。我们常说的0603封装就是指英制代码。另一种是米制代码，也由4位数字表示，其单位为毫米。下表列出贴片电阻封装英制和公制的关系及详细的尺寸：

一、零件规格: (a)、零件规格即零件的外形尺寸，SMT发展至今，业界为方便作业，已经形成了一个标准零件系列，各家零件供货商皆是按这一标准制造。 标准零件之尺寸规格有英制与公制两种表示方法，如下表 英制表示法1206 0805 0603 0402 公制表示法3216 2125 1608 1005 含义 L:1.2inch(3.2mm)W:0.6inch(1.6mm) L:0.8inch(2.0mm)W:0.5inch(1.25mm) L:0.6inch(1.6mm)W:0.3inch(0.8mm) L:0.4inch(1.0mm)W:0.2inch(0.5mm) 注： a、L（Length）：长度；W（Width）：宽度；inch：英寸 b、1inch=25.4mm （b）、在(1)中未提及零件的厚度，在这一点上因零件不同而有所差异，在生产时应以实际量测为准。 （c）、以上所讲的主要是针对电子产品中用量最大的电阻（排阻）和电容（排容），其它如电感、二极管、晶体管等等因用量较小，且形状也多种多样，在此不作讨论。 （d）、SMT发展至今，随着电子产品集成度的不断提高，标准零件逐步向微型化发展，如今最小的标准零件已经到了0201。 二、常用元件封装 1）电阻： 最为常见的有0805、0603两类，不同的是，它可以以排阻的身份出现，四位、八位都有，具体封装样式可参照MD16仿真版，也可以到设计所内部PCB库查询。