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Frequency Measurement of Musical Instrument Strings Using Piezoelectric Transducers

Abstract and Figures

The use of a piezoelectric transducer to monitor the tuning of a musical instrument string has been investigated. It has been shown that the transverse resonance frequencies of the string can be identified by electrical measurements on a low-cost actuator/sensor, sufficiently discreetly to be done during a performance. This frequency measurement approach can be used as the basis for a tuning control mechanism to maintain a musical instrument string at the required pitch, without it having to be plucked or played. Such a system would be of direct benefit to harp players in particular, who have no other means to adjust a mistuned string during a performance. Some of the practical issues and implications of adding such a tuning control system to the harp are considered.
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Frequency Measurement of Musical Instrument
Strings Using Piezoelectric Transducers
Nicolas Lynch-Aird 1,*,ID and Jim Woodhouse 2,ID
1The Old Forge, Burnt House Lane, Battisford, Suolk IP14 2ND, UK
2Department of Engineering, University of Cambridge, Trumpington St, Cambridge CB2 1PZ, UK;
The authors contributed equally to this work.
Received: 28 November 2017; Accepted: 9 January 2018; Published: 13 January 2018
Abstract: The use of a piezoelectric transducer to monitor the tuning of a musical instrument string
has been investigated. It has been shown that the transverse resonance frequencies of the string can
be identified by electrical measurements on a low-cost actuator/sensor, suciently discreetly to be
done during a performance. This frequency measurement approach can be used as the basis for a
tuning control mechanism to maintain a musical instrument string at the required pitch, without it
having to be plucked or played. Such a system would be of direct benefit to harp players in particular,
who have no other means to adjust a mistuned string during a performance. Some of the practical
issues and implications of adding such a tuning control system to the harp are considered.
Keywords: string; tuning; harp; piezoelectric; resonance
1. Introduction
Stringed instruments are sensitive to temperature and humidity changes. In particular, plain
gut or nylon strings on a harp can be suciently sensitive to environmental changes as to become
noticeably out of tune during the course of a performance. For instruments like the violin, the player
can make corrections by changing their fingering. For open-string instruments such as the harp,
though, the player can only wait for a break in the music to retune an errant string.
Automated string tuning systems have been developed [
], but these generally require the
instrument to be switched into a specific tuning mode, with the string being plucked or struck in order
to generate a tone that can be compared with the required musical pitch. A motor is then used to adjust
the string tension as necessary. Even if such a system were adapted for use during a performance,
it would not help with the situation where the string is already noticeably out of tune when it next
comes to be played. Furthermore, any such system will always require multiple tuning iterations for
each string because the adjustment scaling factor, used to convert the measured tuning error into the
required string winding adjustment, will never be known exactly.
A study of the mechanical properties of centreless-ground (“rectified”) nylon harp strings has
shown that holding such strings at a constant tension may be enough to keep them acceptably in
tune [
]. Many players, however, prefer natural gut strings, which are much more sensitive to changes
in humidity and can change in unpredictable ways: on the same instrument, a given gut string may go
sharp while an adjacent string goes flat. Ideally, a tuning control scheme needs some way of measuring
the pitch of a string without the need for that string to be plucked and without making an audible
noise that would disrupt a performance going on around the instrument. Tuning corrections could
then be made repeatedly, and discreetly, maintaining the string at the correct pitch and ready for use.
The simplest way to monitor an individual string, without interfering with its playing properties,
is to make the necessary measurements at one or other of the string’s terminations. Some kind of
Vibration 2018,1, 2; doi:10.3390/vibration1010002
Vibration 2018,1, 2 2 of 17
actuator would be used to apply a low-level mechanical input, and some kind of sensor would measure
the response. Ideally, these functions would be combined in the same device, which would also be
inexpensive and robust. Such a system is described and tested in this paper, based on a low-cost
piezoelectric disc transducer. Electric and electro-acoustic harps often use piezoelectric pickups,
with each string passing over a separate pickup at the point where the string meets the soundboard.
However, these pickups have a dierent structure from the transducer used here, and it will be
demonstrated in Section 4.7 that they are less eective for the purposes of monitoring tuning in the
manner explored here.
As well as exploring the ability to measure the string pitch, the study also investigates the eects
of varying the string stress and the use of dierent drive levels. An approximate theoretical model
is developed and compared with measurements; and the performance of a possible tuning system
is demonstrated. Finally, some consideration is also given to other possible pitch measurement
approaches based on the use of piezoelectric transducers.
2. Methods and Materials
2.1. Test Rig
A simple test rig (Figure 1) was constructed on which a string could be mounted approximately
13 mm above a wooden baseboard. A motorised string winder was mounted at one end of the string,
allowing precise string length adjustments to be made when required. From the winding shaft, the
string passed a fixed vertical bridge pin 5.5 mm in diameter. At its other end the string passed over a
small bridge resting on the piezoelectric transducer, and then sloped down to a clamping point on the
baseboard. The bridge consisted of a semi-cylindrical steel spacer, 4 mm in length, holding the string
about 2.5 mm above the transducer disc. The mass of the spacer was about 0.2 g. The termination
points at each end of the vibrating length of string (the vertical bridge pin and the horizontal spacer)
were set at 90
to each other so that the eective vibrating length of the string was not aected by the
polarisation direction of the transverse vibration [3,4].
string winder
Wooden baseboard
bridge pin
Tail section
End clamp
Figure 1. A schematic representation of the test rig.
The string used for these tests was a centreless-ground nylon string from the Bowbrand
“Pedal Nylon” range, with an unstretched diameter of 0.84 mm and an unstretched linear density of
0.593 g/m. The string vibrating length
between the vertical bridge pin and the piezoelectric transducer
was 223 mm. The length of the string tail section, from the piezoelectric transducer down to the end
clamp attached to the baseboard, was 66 mm, with a break angle of about 3°.
For the main tests to be described, the transducer used was a low-cost disc unit marketed as a
piezoelectric sounder (Maplin QY13P [
]). This consists of a circular metal base plate 27 mm in diameter
and 0.3 mm thick, with a piezoelectric disc 20 mm in diameter and 0.2 mm thick bonded to one face.
The base plate forms one electrode, while a thin metal coating on the other side of the piezoelectric
disc provides the second electrode. This type of transducer is designed so that the piezoelectric layer
contracts or expands radially in response to an applied electrical voltage, causing the base plate to
bend in an axisymmetric manner: it ‘domes’ in and out, similar to the fundamental mode of vibration
of a circular diaphragm. To allow this motion to occur unimpeded, the transducer was supported
Vibration 2018,1, 2 3 of 17
on an annular cup washer with a cavity under most of the base plate. Direct measurements with a
laser-Doppler vibrometer confirmed that the transducer did indeed move in the expected direction,
normal to the plane of the base plate [
]. Figure 2shows the piezoelectric transducer mounted on its
annular washer, with the string running over the semi-cylindrical spacer positioned in the centre of the
top electrode.
Figure 2.
The piezoelectric transducer sitting on top of its circular support. The string is held away from
the transducer and support with a semi-cylindrical steel spacer, with a mass of about 0.2 g. The main
string section enters from the left of the picture, while the tail section and end clamp are to the right of
the transducer assembly.
Following an initial series of investigations, the test rig was left for a period of about six months
prior to taking the set of measurements reported here. This had the advantage of allowing plenty of
time for the string to settle. While variations in the string tension and frequency could be expected
to result from changes in the ambient temperature [
], the string was (at least for the purpose of
these tests) no longer creeping. To take advantage of this stability, the test sequence was organised to
complete as many measurements as possible before making any tuning adjustments.
2.2. Electrical Measurements
In operation, a sinusoidal voltage was applied to the transducer, producing normal motion
which drove the string into vibration. That vibration resulted in a reaction force at the bridge,
causing additional axisymmetric bending motion in the base plate of the transducer (which acts as a
spring). This in turn produced strain in the piezoelectric layer, which has the eect of modulating
the capacitance of the piezoelectric transducer. This modified capacitance can be monitored via an
electrical measurement.
The electrical impedance
of the piezoelectric transducer was measured as a function of
frequency using the bridge circuit shown in Figure 3. In order to select the reference capacitance
Cre f
, some idea of the transducer capacitance was required. Initial measurements were made using
a home-built capacitance meter which applied positive and negative voltage steps, via a reference
resistor, and measured the rise and fall times of the voltage across the capacitance. This had a test
cycle frequency of about 42 Hz, so the values obtained should be close to the DC capacitance for
the piezoelectric transducer. Using this approach gave a capacitance of around 23–24 nF when the
transducer was removed from the rig and able to move freely. The reference impedance for the other
arm of the bridge was therefore selected to be a capacitor of about the same size, with a nominal value
of 22 nF, and a measured value (using the same capacitance meter) of 26 nF.
For the intended test frequency range the impedance of both the test and reference capacitors was
expected to be in the region of 10 k
. The resistive elements of the bridge,
, were chosen
accordingly. To allow easy exploration of the eects of variations in the bridge balance, these two
Vibration 2018,1, 2 4 of 17
components were provided using a single 20 k
linear potentiometer. The instrumentation amplifier
used for measuring the voltage dierence across the bridge (Texas Instruments INA121P) was chosen
for its very high input impedance (10
) and the ability to set its gain using a single external resistor.
A bias current path to ground was required for each input, and two 10 M
Rre f
were included for that purpose. A second instrumentation amplifier was connected directly across the
piezoelectric transducer for use in recording manual pluck responses from the string. The plucks were
made using a plastic guitar plectrum.
Figure 3.
The impedance measurement bridge circuit.
Cre f
=26 nF;
Rre f
=10 M
were provided by a 20 k
potentiometer which could be adjusted to change the bridge balance.
The instrumentation amplifier used was an INA121P, set for unit gain.
A two-channel digital oscilloscope and function generator (Velleman PCSGU250) was used
both to provide the input signal to the impedance bridge, and to measure and record the output.
While having limited input resolution and sampling duration, this low-cost unit had the advantage
of being straightforward to operate under software control. One oscilloscope channel was used to
measure the bridge output, while the other channel was switched between recording the output
of the signal generator, when measuring the transducer impedance, and the output of the second
instrumentation amplifier, when recording pluck responses.
As will be explained in Section 4.3, it is beneficial to measure the bridge output relative to the input
signal in complex form. Sine functions were fitted to the recorded waveforms for the bridge input and
output using a least-squared-error approach: a function
of known frequency
can be fitted as
y=A+Psin ωt+Qcos ωt
P=Bcos C
Q=Bsin C
. The magnitudes and
phases of the fitted functions were then combined to give the complex output.
Tests were also run using a simple voltage divider configuration, with just the reference and
test impedances connected in series. In this configuration, however, the overall testing sensitivity
is seriously limited by any gain dierences in the instrumentation amplifiers, oscilloscope input
circuits or analogue-to-digital converters used to measure the voltages across the two impedances.
The bridge circuit, while being somewhat more complex to analyse, was found to give significantly
better performance.
Vibration 2018,1, 2 5 of 17
3. Theoretical Background
3.1. Mechanical System
The mechanical system is shown in idealised form in Figure 4. A perfectly flexible string of length
, tension
and mass per unit length
is rigidly fixed at one end (at
0), and terminated at the
other end (
) by a spring of stiness
, representing the transducer disc. A voltage
to the transducer produces a force
on the string, where
is a constant and
is angular
frequency [7]. For this harmonic response problem, the string displacement must take the form:
u(x,ω) = Asin ωx
is a constant and
is the wave speed. Force balance at the junction of the string and
the spring then requires:
x(L,ω) + Ku(L,ω) = TA ω
ccos ωL
c+KA sin ωL
c. (2)
Thus, the receptance of the coupled string and spring is given by:
R(ω) = Y
Z0ωcot ωL
where Y(ω) = u(L,ω)and Z0=pTµis the string’s characteristic impedance.
x = 0 x = L
Figure 4. Sketch of the mechanical model, with a stretched string terminated in a spring representing
the transducer stiness.
The stiness
will turn out not to play a very important role in the measurements, but it is useful
to have an idea of its value. This can be estimated from a standard formula for the central displacement
of a circular plate with simply-supported edges, in response to a concentrated load near the centre
([8], see Table 24):
for a plate of radius
and thickness
, made of material with Young’s modulus
and Poisson’s ratio
Using the dimensions given above for the transducer base plate, and typical values
100 GPa and
0.33 for brass, the result is
30 kN/m. The actual stiness will be higher, for two reasons: first,
there will be a contribution from the piezoelectric layer; second, the eective diameter will be smaller
than the overall diameter because of the support details of the annular washer, and those supports
may also provide boundary conditions a little stier than idealised hinged supports.
Various additional elements were considered for inclusion in the mechanical model, including
the mass of the spacer between the string and the piezoelectric transducer, and the elastic stiness
of the tail section of the string pulling the string against the transducer. The only addition found to
make a significant contribution was internal damping of the mechanical system, which is needed to
reduce resonance peaks to finite magnitudes. Modal frequencies and associated damping factors of the
string on the test rig can be readily measured from pluck tests. To incorporate these into the theoretical
Vibration 2018,1, 2 6 of 17
model, the first few poles of the receptance
can be calculated, with their corresponding residues.
Newton-Raphson iteration starting near the known resonance frequencies of the rigidly mounted
string (i.e., with K→ ∞) was used to compute these poles by finding roots ωnof the equation:
Z0ωn+Ktan ωnL
c=0. (5)
Using l’Hôpital’s rule, the corresponding residues pnare given by:
pn=Z0cot ωnL
. (6)
A damped version of
can then be found by using the same residues, but replacing the real
pole frequencies ωnby complex values incorporating the measured Q factors Qn:
ωn=ωn 1+i
2Qn!. (7)
For the simplest case with just the first pole, the result takes the form:
where the initial constant term is the DC value R(0). More poles could be added in if desired.
3.2. Electrical System
The displacement
is proportional to the induced strain in the piezoelectric layer, so the
transducer capacitance will satisfy:
is the baseline value,
is the voltage across the device, and
is a constant.
could be
replaced by
to allow for internal damping. The analysis of the bridge circuit of Figure 3in the
presence of a time-varying capacitance requires a little care. For simplicity the resistors
Rre f
will be omitted from this calculation, because their eect is negligible.
The directly measured bridge output
is the voltage dierence across the bridge
normalised by the drive voltage
. Throughout this paper, “bridge output” will refer to this normalised
voltage dierence. If the current in the left-hand branch of the bridge is I1, then:
dt C(t)V1(t). (10)
were purely sinusoidal, as one would at first think, the product
would only
involve the varying capacitance through a nonlinear term which would appear at frequency 2
However, empirically it is found that the output is dominated by sinusoidal variation at frequency
The missing ingredient is a feature of the piezoelectric layer which has not yet been mentioned. As a
result of the poling process during manufacture of the device, a remanent electric field is “frozen” into
the interior of the layer [
]. The eect can be represented approximately by including a DC component
of voltage V0in addition to the sinusoidal component. A linearised approximation is then possible:
dt C(t)V1(t)d
dt n(C0+C0eiωt)(V0+V1eiωt)oiω[C0V1+V0C0]eiωt(11)
C0=βY=αβV1R(ω). (12)
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Equation (10) can then be solved for
in this linearised approximation for harmonic response,
and combined with the corresponding linear result for the right-hand branch of the bridge to yield:
B(ω) = V1V2
1+iωRA[C0+αβV0R(ω)] 1
1+iωRBCre f
1+iωRBCre f
where the binomial expansion has been used in the final approximation, on the assumption that the
eect of the term involving
is relatively small. The first two terms on the right-hand side of
Equation (14) cancel if the bridge is perfectly balanced, leaving the third term which is proportional
to the receptance
(which can be replaced by
to represent the eect of damping). Thus,
a measurement of
with a well-balanced bridge would be expected to show a peak near each resonance
of the string, exactly as hoped in the context of designing an automated tuning system. Note that the
three calibration constants
appear only via their combined product, so that there is only a
single value to be identified from measurements on the real transducer. The response is modified if the
bridge is out of balance, as will be explored in Section 4.3.
4. Results
4.1. Initial Characterisation of the Unloaded Piezoelectric Transducer
The first test was an impedance scan of the unloaded piezoelectric transducer, removed from the rig
and without load from the string. Figure 5a shows the bridge output response as a function of frequency.
The values of RAand RBwere measured using a digital multimeter as 9.38 and 10.94 k, respectively.
Figure 5.
) The normalised bridge output response of the unloaded piezoelectric transducer;
) the absolute value of the corresponding impedance of the piezoelectric transducer, assuming it
could be treated as predominantly reactive, plotted on log-log scales; (
) the corresponding capacitance
response of the unloaded piezoelectric transducer. All three plots cover the range 0.1–6 kHz.
Vibration 2018,1, 2 8 of 17
If the transducer impedance is purely reactive, its value can be obtained in a straightforward
manner from the magnitude of the bridge output. Figure 5b shows the absolute value of the derived
reactance response, plotted on a log-log scale. As expected the low-frequency part of this plot, up to
about 2 kHz, shows a straight-line response with gradient
1, indicating that the transducer can
indeed be treated as a capacitance at low frequencies. Figure 5c shows the corresponding capacitance
response on a linear frequency scale: this should only be considered valid over the low-frequency
region. The average capacitance over the frequency range 200 Hz–1 kHz was 24.4 nF, matching the
value obtained previously with the capacitance meter.
Within the electrical engineering community, the behaviour near the first mechanical resonance of
a simple piezoelectric transducer of the type used in this study is usually modelled by an equivalent
circuit consisting of a series LCR circuit in parallel with a capacitance [
]. Presumably this linear circuit
analysis relies on the same DC approximation of the remanent field as used above, although this is
rarely made explicit in the recent literature. The values of the capacitance and inductance in the LCR
arm can be obtained from the series and parallel resonance frequencies, which lie close to the points of
minimum and maximum impedance respectively. The dashed blue line in Figure 5a shows the result
of fitting such a model to the measured bridge output response. The fit obtained is excellent, with only
a secondary peak at about 4.9 kHz being unexplained by this simple model. The series resonance
frequency obtained was 4.3 kHz, which was reassuringly close to the manufacturer’s nominal value of
4.2 ±0.5 kHz.
It is important to the intended sensing mechanism that the low-frequency capacitance of the
transducer, in the region of the fundamental frequency of the string, but below the mechanical
resonance of the transducer, should vary with the applied load. To explore this eect, the transducer
was mounted on the test rig in the normal way (Figures 1and 2), and the applied load was varied by
altering the string tension. In each case the capacitance was derived from the low-frequency bridge
output response. Figure 6shows two sets of results obtained, on dierent days (blue and cyan plots).
The vertical shift between the two cases is due to sensitivity of the transducer capacitance to changes in
the ambient temperature, probably a consequence of thermal sensitivity in the dielectric constant of the
piezoelectric material [
]. An important implication for any practical application is that an impedance
bridge of the type used here will not remain balanced as the ambient temperature varies, with a net
contribution from the first two terms of Equation (14) appearing in the bridge output response.
Figure 6.
Variation in the transducer capacitance with the applied static load. For the blue and cyan
plots the transducer was mounted on the test rig in the normal way (Figure 2), and the load was varied
by altering the string tension. For the green plot, the transducer was instead mounted on a wire loop
support, which enabled the transducer to be turned over to measure the eect of a reversed load.
Vibration 2018,1, 2 9 of 17
With the transducer mounted in the normal way, the load could only be applied in one direction.
To obtain the green plot in Figure 6, the support arrangement was temporarily changed to use a
horseshoe-shaped loop of wire instead of the cup washer, which allowed the transducer to be turned
over without the support interfering with its leads. Tests were run with the transducer in both its
normal orientation (electrode on top) and upside down. In this case, a third point was provided by the
capacitance of the unloaded transducer, as given above. All three sets of responses show a clear linear
trend, and the fitted lines have consistent gradients. The rate of variation of capacitance with load
is perhaps surprisingly large, which results in high sensitivity for the dynamic measurements to be
described shortly.
4.2. Loaded Transducer Response and String Resonances
The piezoelectric transducer was replaced in its normal position on the test rig, as shown in
Figure 2, and an impedance scan was recorded. A manual string pluck response was also recorded for
comparison. Figure 7shows the bridge output (black) together with the FFT of the pluck response
(red). The FFT has been scaled to suit the vertical scale range of the bridge output response. The bridge
output response from the unloaded transducer (dotted black) is included for reference. Spikes in the
bridge response can be clearly seen near the string fundamental (834.6 Hz) and the first few overtones.
The spike at about 2.9 kHz was the fundamental for the string tail section (Figures 1and 2). Comparing
the loaded and unloaded responses, it can be seen that the mechanical resonance of the transducer
had fallen significantly. The low-frequency capacitance had also fallen, relative to the unloaded case,
to 23.3 nF; consistent with Figure 6. Note that the frequency region of interest, around the string
fundamental, was still well below the transducer resonance; so the transducer could be treated as a
capacitance for the purposes of the model developed in Section 3.
Figure 7.
The bridge output for the piezoelectric transducer loaded by the string (black solid line).
The FFT of a manual pluck response is shown for comparison (red solid line). Spikes in the impedance
response can be seen near the first few string resonances. The spike at about 2.9 kHz corresponds to
the fundamental resonance of the string tail section. The bridge output response for the unloaded
transducer is included for reference (dotted black line).
4.3. Variations in Impedance Bridge Balance
When viewed at the plotting scale of Figure 7, it appears that the bridge output shows a peak at
the string resonance frequency. However, closer examination confirms the prediction of Equation (14):
the precise peak frequencies in measurements like the solid black curve in Figure 7are influenced by
Vibration 2018,1, 2 10 of 17
the state of balance of the bridge circuit. Using a value for the transducer capacitance of
23.3 nF,
Cre f =
26 nF, suggested that the values of
for the impedance bridge to be balanced
should be 10.5 k
and 9.5 k
respectively. The initial values of
9.38 k
10.94 k
therefore corresponded to the bridge being somewhat out of balance. For the bridge to be out of
balance by about the same amount, but in the opposite direction should have required
11.74 k
and RB=8.26 k.
Figure 8shows the bridge output responses obtained around the string fundamental (834.6 Hz)
with approximately these three resistance ratios, plus another two for approximately twice the degree
of imbalance in either direction. For the green response it appears that the bridge setting was indeed
close to balanced, with a single response peak at the string resonance, as expected from Equation (14).
The magenta response corresponded to the bridge balance condition for the previous tests, with the
response dipping prior to rising to a distinct peak as the frequency increased through the string
resonance. The red response, with the bridge further out of balance in the same direction, has a similar
shape to the magenta response, but with an increase in its vertical oset. Changing the bridge to be out
of balance in the opposite direction (cyan and blue) gave responses with a very similar shape, but with
the dip now following the peak as the frequency increased. In the four unbalanced cases it can be seen
that the size of the vertical oset has increased with the degree of imbalance in the bridge; as expected
from the first two terms of Equation (14).
Figure 8.
The output response around the string fundamental frequency (834.6 Hz) with dierent
settings for the resistive arms of the impedance bridge. When the bridge was close to balance (green),
the output response had a single peak close to the string resonance. When the bridge was suciently
out of balance, the response had a distinct peak and trough, with the string resonance lying somewhere
between these two points. Changing the direction of the bridge imbalance reversed the order of the
peak and trough, but not the overall form of the response. The dashed lines show the predicted
responses from Equation (14).
The dashed lines in Figure 8plot the responses predicted by the simple model of Equation (14),
replaced by
and using just the first pole, as in Equation (8). It was found that
the first term on the right-hand side of Equation (8), which for musical instrument strings will be
dominated by the transducer stiness
, actually made only a very small contribution, which could be
allowed for by a small adjustment in the low-frequency transducer capacitance
. It was expected that
a temperature-related adjustment would be needed anyway in the value of
, since the measurements
presented in this section were made on a dierent day to the measurements presented in Section 4.2.
Taking this approach, of ignoring the first term on the right-hand side of Equation (8), it was then only
necessary to fit a value for the combined scale factor αβV0p1.
Vibration 2018,1, 2 11 of 17
The value for the Q factor at the string fundamental was obtained by analysing a recorded pluck
response, as described in Section 3, giving
655. The value of
was set to 23.5 nF instead of the
value 23.3 nF obtained from the loaded impedance response (Section 4.2), while the combined scale
factor used was
. As can be seen from Figure 8, the match is excellent, suggesting
that the simple model given in Equation (14) is capturing the key behaviour rather well.
Referring again to Equation (14), the net eect of the first two terms should be to add a complex
oset that varies only slowly with frequency compared to the rapid variation associated with the
resonance peak. This suggests a possible approach to “balancing” the bridge in software: if an average
value of the complex oset can be estimated, by examining the bridge output response away from the
string resonance, the measured bridge output can be corrected by subtraction of this constant oset.
The result should be a clear peak at the string resonance.
Figure 9demonstrates this approach. Figure 9a is a repeat of the measured bridge output
magnitude responses from Figure 8, while Figure 9b shows the evolution with frequency of the complex
values of these responses. It can be seen that it is only necessary to move a few Hz away from the string
resonance for the associated response deviations to have become relatively insignificant, compared to
the slowly-varying background oset, with the latter varying with the bridge balance setting. Complex
oset corrections, shown as dashed lines in Figure 9b, were calculated by taking the averages of the real
and imaginary responses over the 1 Hz intervals at either end of the measured responses, i.e., between
about 14 and 15 Hz either side of the string resonance. Figure 9c shows the results of applying these
balance corrections. The shape and amplitude of the corrected responses match very closely. In fact,
this approach was sensitive enough to identify small variations in the string resonance, probably due to
temperature changes between successive measurements while the bridge balance was being adjusted.
These osets, relative to the green (near-balanced) response, and in the order that the responses are
listed in Figure 8, amounted to 0.5, 0.1, 0.0, 0.0, and 0.1 Hz respectively. Corresponding corrections
have been added to the frequency values in Figures 8and 9.
Figure 9.
) A repeat of the measured magnitude responses from Figure 8; (
) the real and imaginary
parts of the same bridge output responses; (
) the adjusted magnitude responses, after adding a balance
correction oset.
Vibration 2018,1, 2 12 of 17
4.4. Variations in Drive Level
Most of the tests reported here were run with drive waveform levels in the range 6–10 V
produce clear response plots. At these drive levels, however, the tones generated by the piezoelectric
transducer and string were quite audible, especially at the fundamental string frequency. Figure 10
shows the bridge output responses obtained with lower drive levels, reducing the drive amplitude by
up to a factor of 30. At the lowest levels the test tones could only just be heard at the string fundamental.
On the other hand, the test string was not attached to a soundboard which, by design, would be
expected to amplify any sound coming from the string. Encouragingly, the bridge output responses
obtained still showed a clear variation in amplitude, with a distinct dip and peak: the shape of the
bridge output response did not change. These results also demonstrate the linearity of the overall
measurement approach. With careful design of the bridge and sensing electronics and wiring, and also
of the piezoelectric transducer, the drive levels could almost certainly be reduced further. The drive
level could also be varied adaptively during the frequency scan, reducing it as the measured response
passes through its peak.
Figure 10. The bridge output responses for dierent input amplitudes.
4.5. Variations in String Stress
The longitudinal stress in the non-metallic (nylon or gut) strings of a concert harp range from
about 50 MPa for the lowest notes up to about 270 MPa for the top notes [
]. To explore what eect the
string stress might have on the ability to measure its tuning by the proposed method, the test string
was retuned to a number of fundamental frequencies, corresponding approximately to stress levels
from about 30–210 MPa in 30 MPa steps. Such large tuning changes inevitably caused the string to
start creeping again, but, since the aim was to investigate the general nature of the corresponding
bridge responses rather than their precise details, some creeping in the string resonance during the
impedance scans did not present a problem.
Figure 11 shows the results obtained. In every case, a clear spike in the balanced bridge output
was seen. The magnitude of the balanced output peak was smaller at low stress levels, and increased
fairly steadily with the fundamental frequency. This suggests that this impedance scanning technique
could be used as the basis for measuring the frequency of unplucked strings across the full range of
gut or nylon strings on the harp.
Vibration 2018,1, 2 13 of 17
Figure 11.
The balanced bridge output responses obtained around the string fundamental frequencies
for a range of stress levels.
4.6. Tuning Performance
Based on these encouraging results of the proposed measurement method, a tuning algorithm
was written which used impedance scanning to determine the string frequency and the necessary
tuning adjustment. Calculating the length of string to be wound onto or othe winding shaft to
make a given tuning adjustment requires knowledge of the string’s Young’s modulus: specifically,
the string’s “tangent modulus” (the stress/strain ratio for small perturbations). For nylon strings this
quantity varies significantly with the string stress, but the previous study of nylon harp strings led to
an expression which can be used to predict the Young’s modulus with sucient accuracy to provide a
stable tuning control algorithm [3].
The tuning algorithm measured the string fundamental frequency using the approach described in
Section 4.3: calculating and applying a complex oset to balance the measured bridge output response,
and then identifying the fundamental string frequency from the peak value. For each tuning iteration,
the algorithm used the measured frequency and the desired target frequency to determine the scanning
range to be used for the next frequency measurement. For a given target frequency
and tuning error
, the scanning range was set to
0, or the symmetrical equivalent
0. This approach was found to be more robust, than just assuming an accuracy margin for
the tuning adjustments, when dealing with a creeping string. For the first iteration, the frequency
measurement was determined from a manual string pluck. This matches the way such a tuning
approach would likely be used in practice, with the instrument initially tuned on the basis of manual
plucking or playing of the strings, before the start of a performance, and with the impedance-based
tuning algorithm being used to maintain the string frequency during the performance.
Figure 12 shows tuning behaviour from two dierent test runs: one without any tuning correction
applied, and one with the tuning algorithm in operation. At the start of the test the frequency of the
string, which had been allowed to settle overnight, was measured as 801 Hz. The string was then tuned
up to 844 Hz (a big enough step to ensure it started creeping again) and its frequency was monitored
using the impedance scanning technique, but without making any corrections. The magenta line in
Figure 12 shows the results obtained. The tuning error is given in cents, there being 100 cents in one
equal-tempered semitone [10]. It is generally accepted that the minimum perceptible change in pitch
of a musical note is never lower than about 3 ¢; the exact value varies with the pitch and complexity of
the sound [
]. For most of the test period the tuning error became increasingly negative, indicating
the string was going flat: the string was creeping, causing its tension and pitch to fall. The red section
Vibration 2018,1, 2 14 of 17
of the response indicates a period when a desk lamp, with a 60 W incandescent bulb, was placed near
the string to provide some local heating. Raising the temperature of a settled nylon string makes it go
sharp [
]. In this case, however, the eect of heating merely oset the underlying creep for a while.
It was also found during the previous study of nylon strings [
] that raising the temperature above the
historic maximum temperature, to which a string had been exposed, would trigger a further episode
of rapid creep. This appears to have happened part way through the period of heating (this test was
run using a new string after the previous string was broken by accidentally over-tightening it).
At the end of the first run the string frequency was re-measured as 830 Hz. The string was then
tuned up to 875 Hz (a similar tuning jump to the first run), and now the tuning control algorithm was
activated. The blue line in Figure 12 shows the results, with the red section again indicating a period
of local heating. With the exception of some short periods, after the initial large tuning change and
at the start and end of the heating period, the control algorithm maintained the string tuning within
2 ¢, and at no time did the tuning error exceed 3.5 ¢. It is also worth noting that for both of these
test runs the input drive level was reduced to 0.7 V
, with no noticeable degradation in performance
(see Section 4.4).
Figure 12.
Tuning variations with and without the tuning control algorithm. In both cases the string
fundamental was increased by about 45 Hz at the start of the test run, causing the string to start creeping.
The red sections indicate periods when local heating was applied by placing a 60 W incandescent bulb
close to the string. The dashed lines are at ±3 ¢.
4.7. Other Measurement Configurations
Some tests were also run using a piezoelectric harp pickup from Camac Harps. This was a
rectangular piezoelectric plate approximately 7
3 mm in size, about 1 mm thick, and with a capacitance
of 440 pF. In common with many other piezoelectric pickups used with musical instruments, this
transducer was not intended to flex in any way: it was presumably designed to generate a voltage
proportional to the force exerted by the string via bulk compression or shear, with very little associated
deformation. This, in conjunction with the pickup’s small size, meant that it was not possible (at least
with the available drive voltages of up to about 30 V
) to obtain a useful impedance scan response
with this transducer.
As an alternative frequency measurement approach, the configuration of the test rig was changed
so that the string passed over the Maplin piezoelectric disc transducer at the end near the motorised
winder, and over the Camac pickup at the other (clamped) end. Tones were injected into the string
using the Maplin transducer, and the responses measured at the other end using the Camac pickup.
Vibration 2018,1, 2 15 of 17
As expected, there was a detectable increase in the output from the Camac pickup when the test
frequency passed through the string resonance, but the overall performance was far inferior to that
obtained using the impedance scanning approach with the Maplin transducer. This suggests that
alternative pitch measurement techniques based on injecting a pulse-train or some other signal at one
end of the string, and measuring the response at the other end, may also perform relatively poorly.
5. Discussion and Conclusions
This study has shown that the transverse resonance frequencies of a stretched flexible string
can be determined from the electrical impedance response of a piezoelectric transducer in contact
with the string at one of its endpoints. The successful measurements were made using a low-cost
circular piezoelectric transducer, 27 mm in diameter, designed to flex its base plate. The transducer
is driven with a sinusoidal waveform at a low level, inducing some barely-audible vibration in the
string. The vibrating string then exerts force back on the transducer, resulting in a modulation of its
capacitance which can be detected using a bridge circuit. By sweeping the sinusoid over a range near
the fundamental resonance of the string, a peak at the resonance frequency can be detected.
The capacitance of a piezoelectric transducer, below its own mechanical resonance, can be expected
to vary noticeably with temperature due to the thermal sensitivity of the dielectric constant of the
piezoelectric material [
]. Consequently, the impedance measurement bridge will usually be operating
somewhat out of balance. The result is to obscure the precise frequency of the string resonance.
However, through the development of a simple explanatory model for the transducer behaviour,
a straightforward approach has been demonstrated for correcting the bridge balance in software.
The string resonance can then be robustly and accurately identified.
It has also been demonstrated that this frequency measurement approach can be used as the basis
for a tuning control mechanism, capable of maintaining a musical instrument string at the required
pitch, without the string having to be plucked or played. An instrument equipped with such a system
would be stabilised against changes in both the strings and in the instrument frame and soundboard.
This would be of direct benefit to musicians, and to harp players in particular, who have no other
means to adjust a mistuned string during a performance.
Tests using a much smaller Camac harp pickup, designed to work with small bulk strains rather
than by flexure, were not able to distinguish the string resonance via the transducer impedance
response. This suggests that a quite dierent transducer design will be required from that normally
used for the pickups in electric and electro-acoustic harps. This need not be a drawback, since it seems
likely that the transducers for auto-tuning purposes would work best if mounted on a base that is as
solid and inert as possible: this would reduce any influence of vibration of the base structure and also
minimise cross-coupling between adjacent strings. This suggests positioning the transducers at the far
end of the string, away from the soundboard.
A brief attempt was also made to measure the string resonance by injecting tones at one end of the
string and measuring the response amplitude at the other end. While the resonance could be identified,
as a peak in the output level, the drive levels required for successful operation were much higher than
with the impedance measurement approach. Such an approach would have the additional drawback
of requiring a separately wired transducer at each end of the string. Most electric or electro-acoustic
harps that have a separate piezoelectric pickup on each string still have multiple pickups connected in
parallel to a common amplifier circuit.
On the pedal or lever harp, there would be a potential issue with the fourchettes or other semitone
mechanisms, mounted along the neck of the instrument. These work by gripping the string, slightly
shortening its vibrating length and hence raising the sounding pitch. A number of possible solutions
may be available depending on how the tuning control system is to be used. In concert performances,
there are often long periods when the harp is not being played. One possibility, therefore, would be for
the player to disengage the semitone controls (pedals or levers) so that the tuning control mechanism
can operate on the open strings until the harp is to be played.
Vibration 2018,1, 2 16 of 17
If it is desired that the tuning control mechanism should be in more continuous use, then a
means will be needed to deactivate it when the string is played. This could perhaps be achieved by
detecting large-amplitude string excitation, or by using sensors to detect when the player’s hands or
fingers are near the strings. A delay would have to be allowed before reactivating the tuning control
to allow the string vibration to decay. The existing mechanical pedal and lever mechanisms could
then be replaced by electro-mechanical servo units, which automatically disengage the fourchettes,
or equivalent, when the tuning control is active. Alternatively, the piezoelectric transducers could be set
to make contact with the string below the fourchettes, and engaged or disengaged under servo control.
The latter option might allow the tuning control system to be retrofitted to an existing instrument.
Although these suggestions all sound rather complicated, it should be borne in mind that the cost of a
full-size concert harp can run into tens of thousands of dollars. While considerable development eort
would be required, all the necessary components are readily available at relatively low cost.
During the course of a performance, the reference pitch of an orchestra can change, but the
harp player has no means to follow such changes. A system capable of making automated tuning
adjustments, however, could also have its reference pitch adjusted (perhaps manually by the player),
causing the whole instrument to be retuned accordingly.
A further problem that may be encountered by any tuning control system operating only on
the string fundamental stems from the fact that, in practice, there are always two “fundamental
frequencies”. Real strings and their associated termination geometry are never perfectly symmetric,
so that the two polarisations of string vibration will always have slightly dierent resonance frequencies
]. On the test rig, eorts were made to minimise the influence of these split peaks; but an automated
tuning control system on a real instrument would need to cope with this phenomenon.
Supplementary Materials:
The following are available online at
the measurement data files, plus a file guide and the Python code for generating the plots.
The authors thank John Durrell and Sohini Kar-Narayan of the University of Cambridge for
advice on the behaviour of piezoelectric materials; Jakez François of Camac Harps for providing samples
of the Camac piezoelectric pickups; Justin Grimwood for assistance with fabricating the string winder
assembly; David Anderson of Accusound for electronic circuits advice and suggestions; and Matteo Frigo
and Steven G. Johnson for making their FFTW algorithm freely available ( The plots were prepared
using the Matplotlib graphics package [12].
Author Contributions:
Nicolas Lynch-Aird conceived of, designed and performed the experiments. Both authors
contributed equally to analysis of the results and writing the paper.
Conflicts of Interest: The authors declare no conflict of interest.
The following abbreviations are used in this manuscript:
DC Direct Current
FFT Fast Fourier Transform
LCR Inductance Capacitance Resistance
Adams, C. Device and Method for Automatically Tuning a Stringed Instrument, Particularly a Guitar.
U.S. Patent 20080282869, 20 November 2008.
Tronical. How It Works? Available online: (accessed on
16 December 2017).
3. Lynch-Aird, N.J.; Woodhouse, J. Mechanical properties of nylon harp strings. Materials 2017,10, 497.
Woodhouse, J. Plucked guitar transients: Comparison of measurements and synthesis. Acta Acust.
United Acust. 2004,90, 945–965.
Maplin. 3V Ceramic Piezo Transducer 27/4.2. Available online:
piezo-transducer-2742-qy13p (accessed on 16 December 2017).
Vibration 2018,1, 2 17 of 17
6. Green, T. Mechanical Behaviour of Nylon Musical Strings. Master’s Dissertation, Cambridge University,
Cambridge, UK, 2015.
Waanders, J.W. Piezoelectric Ceramics: Properties and Applications; Philips Components: Eindhoven,
The Netherlands, 1991.
Young, W.C.; Budynas, R.G. Roark’s Formulas for Stress and Strain, 7th ed.; McGraw-Hill Professional:
New York, NY, USA, 2002.
Waltham, C. Harp. In The Science of String Instruments; Rossing, T.D., Ed.; Springer: New York, NY, USA,
2010; Chapter 9.
10. Benson, D.J. Music: A Mathematical Oering; Cambridge University Press: Cambridge, UK, 2007; p. 166.
Kollmeier, B.; Brand, T.; Meyer, B. Perception of speech and sound. In Springer Handbook of Speech Processing;
Benesty, J., Sondhi, M.M., Huang, Y., Eds.; Springer: Berlin/Heidelberg, Germany, 2008; Chapter 4, p. 65.
12. Hunter, J.D. Matplotlib: A 2D graphics environment. Comput. Sci. Eng. 2007,9, 90–95.
2018 by the authors. Licensee MDPI, Basel, Switzerland. This article is an open access
article distributed under the terms and conditions of the Creative Commons Attribution
(CC BY) license (

Supplementary resource (1)

For years, friction-induced vibration and noise (FIVN) has puzzled many researchers in academia and industry. Several mechanisms have been proposed for explaining its occurrence and quantifying its frequencies, notably for automotive brake squeal, clutch squeal, and even rail corrugation. However, due to the complex and complicated nature of FIVN, there is not yet one fundamental mechanism that can explain all phenomena of FIVN. Based on experimental results obtained on a simple test structure and corresponding numerical validation using both complex eigenvalue analysis (CEA) and transient dynamic analysis (TDA), this study attempts to propose a new fundamental mechanism for FIVN, which is the repeated cycles of partial detachment and then reattachment of the contact surfaces. Since friction is ubiquitous and FIVN is very common, the insight into FIVN reported in this paper is highly significant and will help establish effective means to control FIVN in engineering and daily life.
Full-text available
The long-term mechanical behaviour of a number of fluorocarbon and gut harp strings has been examined, and the results compared with a previous study of rectified nylon strings. The stretching behaviour of the three materials was studied via different measures of the Young’s modulus; with test time scales on the order of weeks, minutes, and milliseconds. The strings were subjected to cyclic variations in temperature, enabling various aspects of their thermal behaviour to be investigated. The effects of humidity changes on gut strings were also examined. The behaviour of the fluorocarbon strings was found to be similar in many ways to that of the nylon strings, despite their different chemical formulation and significantly higher density. In particular, the faster measures of Young’s modulus were found to show an almost identical strong variation with the applied stress; while the thermal behaviour of both materials was largely determined by the balance between opposing effects associated with thermal contraction and thermal variations in the Young’s modulus. The gut strings showed some similarities of behaviour to the synthetic materials, but also major differences. All three measures of the Young’s modulus remained constant as the applied stress was increased. The gut strings were far more sensitive to changes in humidity than the synthetic materials, although some of the results, especially the thermal tuning sensitivity of the strings when held at constant length, displayed remarkable stability under changing humidity. The observed behaviour suggests very strongly that there is significant coupling between humidity-related changes in the linear density of a gut string and complementary changes in its tension.
Full-text available
Monofilament nylon strings with a range of diameters, commercially marketed as harp strings, have been tested to establish their long-term mechanical properties. Once a string had settled into a desired stress state, the Young’s modulus was measured by a variety of methods that probe different time-scales. The modulus was found to be a strong function of testing frequency and also a strong function of stress. Strings were also subjected to cyclical variations of temperature, allowing various thermal properties to be measured: the coefficient of linear thermal expansion and the thermal sensitivities of tuning, Young’s modulus and density. The results revealed that the particular strings tested are divided into two groups with very different properties: stress-strain behaviour differing by a factor of two and some parametric sensitivities even having the opposite sign. Within each group, correlation studies allowed simple functional fits to be found to the key properties, which have the potential to be used in automated tuning systems for harp strings.
Full-text available
Monofilament nylon strings with a range of diameters, commercially marketed as harp strings, have been tested to establish their long-term mechanical properties. Once a string had settled into a desired stress state, the Young's modulus was measured by a variety of methods that probe different time-scales. The modulus was found to be a strong function of testing frequency and also a strong function of stress. Strings were also subjected to cyclical variations of temperature, allowing various thermal properties to be measured: the coeffcient of linear thermal expansion and the thermal sensitivities of tuning, Young's modulus and density. The results revealed that the particular strings tested are divided into two groups with very different properties: stress-strain behaviour differing by a factor of two and some parametric sensitivities even having the opposite sign. Within each group, correlation studies allowed simple functional fits to be found to the key properties, which have the potential to be used in automated tuning systems for harp strings.
Full-text available
The predictions of a calibrated synthesis model for the coupled string/body vibration of a guitar are compared with measurements. A reasonably good match is demonstrated at low frequencies, except that some individual decay rates of "string modes" are hard to match correctly. It is shown to be essential to use an accurate damping model for the string: the intrinsic damping of the classical guitar strings used here is found to vary by an order of magnitude in the audible frequency range. At higher frequencies, two phenomena not included in the synthesis model are found to produce disparities between measurement and synthesis. First, the two polarisations of string motion show modal frequencies which are split further apart than can be explained by body coupling. This effect is tentatively attributed to slight rolling on the fret and/or bridge, resulting in different effective lengths for the two polarisations. Second, unexpected extra peaks are seen in the measured spectra, which are frequency-doubled relatives of the expected peaks due to transverse string resonances. The non-linear mechanism creating these additional peaks is presumed to involve excitation of longitudinal string motion by the transverse vibration.
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Device and Method for Automatically Tuning a Stringed Instrument
  • C Adams
Adams, C. Device and Method for Automatically Tuning a Stringed Instrument, Particularly a Guitar. U.S. Patent 20080282869, 20 November 2008.
How It Works? Available online
  • Tronical
Tronical. How It Works? Available online: (accessed on 16 December 2017).