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DOI: 10.1243/09596518JSCE588

2008 222: 757Proceedings of the Institution of Mechanical Engineers, Part I: Journal of Systems and Control EngineeringG Di Rito and R Galatolo

primary flight actuatorsExperimental and theoretical study of the electrical failures in a fault-tolerant direct-drive servovalve for

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Experimental and theoretical study of the electricalfailures in a fault-tolerant direct-drive servovalve forprimary flight actuatorsG Di Rito* and R Galatolo

Department of Aerospace Engineering, University of Pisa, Pisa, Italy

The manuscript was received on 4 March 2008 and was accepted after revision for publication on 8 April 2008.

DOI: 10.1243/09596518JSCE588

Abstract: The current paper deals with the study of the electrical failures in fault-tolerantflight actuators, with particular reference to the short circuits of the servovalve coils. A high-fidelity model of the servovalve of a modern fly-by-wire actuator is developed and validatedthrough experiments, focusing attention on the characterization of the component dynamics incase of partial and total short circuits of the direct-drive motor coils. The servovalve model isthen used to simulate a typical on-ground built-in-test procedure to determine the limitcondition for the detection of a partial short circuit. Finally, once different possiblecombinations of short circuits are injected, the degradation of performances of the wholeactuator is characterized through experiments, and the servovalve model is used to justify thetest results, highlighting and discussing the effects of the failures on the system dynamics.

Keywords: flight control, fault-tolerant actuators, direct-drive servovalve, failure analysis,testing, modelling and simulation

1 INTRODUCTION

Full-authority fly-by-wire flight control systems are

nowadays a standard technology solution for ad-

vanced military aircraft and for civil transport

airliners. To provide the adequate level of safety,

the primary flight actuators are designed to be fault-

tolerant components, so that they must be able

appropriately to operate in the event of one or more

failures. The capability of the system to tolerate the

actuator failures also depends on the flight control

computers, which must be able to detect the

malfunctions by means of both on-ground and in-

flight built-in tests [1]. However, as a result of

inadequate or impracticable monitoring procedures,

a flight actuator can operate with dormant failures,

implying an unexpected degradation of the perfor-

mance and the safety of the aircraft.

When hydraulic flight actuators with direct-drive

servovalve [2, 3] are concerned, dormant failures can

be caused by the short circuits in the servovalve

coils. Actually, as these faults can involve only a part

of the windings, the on-ground built-in-tests cannot

detect the malfunction, causing the aircraft to take

off with a damaged actuator. The problem is that a

partial short circuit could extend to the rest of the

coil during the flight, implying a significant degrada-

tion of the actuator performances.

Although many authors have focused their atten-

tion on the study, the modelling, and the dynamic

characterization of electrical motors in faulty condi-

tions [4–6], scarce information is available about the

modelling and the failure analysis of direct-drive

electrical motors for aerospace servovalves.

In the current paper, a high-fidelity model of the

direct-drive servovalve of a modern fly-by-wire

actuator is developed and validated through experi-

ments, focusing attention on the characterization of

the component dynamics in short circuits of the

motor coils.

*Corresponding author: Department of Aerospace Engineering,

University of Pisa, via Caruso, Pisa, Italy. email: g.dirito@ing.

unipi.it

757

JSCE588 F IMechE 2008 Proc. IMechE Vol. 222 Part I: J. Systems and Control Engineering



2 SYSTEM DESCRIPTION

2.1 Architecture of the fault-tolerant flightactuator

As shown in Fig. 1, the reference actuator has a

tandem cylinder and quadruple electronics, so that

the system is capable of tolerating one hydraulic

failure and up to two electrical failures. Concerning

the direct-drive servovalve, its internal architecture

is schematically reported in Fig. 2, where it can be

noted that the valve spool movement is obtained by

a permanent magnet electric motor with four coils

operating in a flux-summed configuration. When no

current circulates in the motor coils, the armature

(rigidly linked to the valve spool) is centred with

respect to its endstrokes, as in this condition the

centring spring is unloaded and the magnetic fluxes

induced by the permanent magnets provide equal

and opposite forces. When the currents are not zero,

the magnetic flux induced by the coils causes an

unbalanced armature polarization; a magnetic force

is therefore generated, and the resulting spool

movement allows the actuator to control the

hydraulic power.

2.2 Electrical failure modes of the direct-driveelectrical motor

Five basic operational states can be defined for each

of the four motor coils:

Fig. 1 Primary fly-by-wire actuator with tandem cylinder

Fig. 2 Quadruple direct-drive motor for aerospaceservovalves

758 G Di Rito and R Galatolo

Proc. IMechE Vol. 222 Part I: J. Systems and Control Engineering JSCE588 F IMechE 2008



(a) the normal condition, in which the coil voltage

is controlled by the electronic drive, composed

of a servo-amplifier performing a closed-loop

current control, Fig. 3(a);

(b) the partial short circuit, in which the short

circuit extends to a limited number of windings,

while the rest of the coil is still connected with

the electronic drive, Fig. 3(b);

(c) the total short circuit, in which a fault results in

a direct electrical connection between the coil

pins, Fig. 3(c);

(d) the hardover, in which the current feedback is

failed and the servo-amplifier fixes the coil

voltage to its maximum value;

(e) the open circuit, in which the coil wire is broken

and the current circulation is inhibited.

The failure analysis proposed in this paper has been

focused on the short circuits, since (as discussed in

section 4) they can imply significant fault detection

problems as well as a severe degradation of the

actuator dynamic behaviour.

3 MODEL OF THE FAULT-TOLERANT DIRECT-DRIVE SERVOVALVE

Before discussing the simulation details, it must be

outlined that the model of the fault-tolerant direct-

drive servovalve has been developed as a multistate

machine, which is able to simulate the component

dynamics in different working conditions (normal;

with one or more total short circuits; with one or

more partial short circuits). However, for the sake of

simplicity, the model will be described with reference

to an operating condition in which a partial short

circuit is applied to one of the four coils, while the

other three coils are active. This condition allows the

basic simulation problems to be highlighted, and the

resulting model can be easily adapted when additional

(partial or total) short circuits are considered.

The model is composed of three sections: one for

the calculation of the magnetic fluxes, another for

the simulation of the current dynamics, and one for

the simulation of the valve spool dynamics.

3.1 Magnetic fluxes in the direct-drive motor

The simulation of the magnetic fluxes has been

obtained by means of the magnetic circuit theory [7],

using (in the case of one partial short circuit) the

magnetic flux scheme reported in Fig. 4. The resulting

magnetic circuit (Fig. 5) contains five coils, as the partial

short circuit has been simulated by dividing the failed

Fig. 3 Motor coil working condition: (a) normal, (b) partial short circuit, (c) total short circuit

Fig. 4 Scheme of the magnetic flux paths in the direct-drive motor in a partial short circuit failure

Electrical failures in a fault-tolerant direct-drive servovalve 759




coil into two sections: one (involving Nsc windings)

totally shorted, and the other ‘normally’ controlled by

the servo-amplifier, thus reproducing the failure mode

scheme reported in Fig. 3(c).

Each magnet is assumed to be a rare-earth magnet

working in the linear range, is modelled as a

magnetomotive force (Wm) in series with a reluctance

(Rim) that represents the demagnetization character-

istics of the material [8]. Each motor coil is modelled

as a magnetomotive force (Njij, where Nj is the

number of coil windings and ij is the coil current in

the jth coil) in series with the reluctances of the air

gaps between one coil and another (Rca for the axial

path and Rcr for the radial path). Finally, the air gaps

between the armature and the case are represented by

the reluctance of the armature radial clearance (Rrc)

and by the reluctances of the axial gaps (Ral and Rar),

which depend on the armature position (xs).

The magnetic field of the motor is thus composed of

ten fluxes: a flux that links all the coils and the axial air

gaps of the armature (Qcn), two fluxes linking the

permanent magnets and the axial air gaps of the

armature (Qml and Qmr), and seven flux leakages,

representing the secondary magnetic paths related to

the permanent magnets (Qmdl and Qmdr) and the motor

coils (Qcd j with j 5 1,…,5).

Once the armature position and the coil currents are

known, the magnetic fluxes can be obtained by invert-

ing equations (1) and (2), and the magnetic force (Fm)

can be calculated by applying the virtual work principle

[7], differentiating the magnetic field energy (Em) with

respect to the armature position, equations (3) and (4)

Wm

Wm

Wm

WmP5j~1

Njij

8>>>>>>>><>>>>>>>>:

9>>>>>>>>=>>>>>>>>;

~

<miz<rcz<ar 0 <mi 0 {<ar

0 <miz<rcz<al 0 <mi <al

<mi 0 <miz<md 0 0

0 <mi 0 <miz<md 0

{<ar <al 0 0 <arz<al

26666664

37777775

Qmr

Qml

Qmdr

Qmdl

Qcn

8>>>>>><>>>>>>:

9>>>>>>=>>>>>>;

ð1Þ

Fig. 5 Equivalent magnetic circuit of the direct-drive motor in a partial short circuit failure





Ni1

Ni2

Ni3

N{Nscð Þi4

Nsci5

8>>>>>><>>>>>>:

9>>>>>>=>>>>>>;

~

<caz<cr {<cr 0 0 0

{<cr <caz2<cr {<cr 0 0

0 {<cr <caz2<cr {<cr 0

0 0 {<cr <caz2<cr {<cr

0 0 0 {<cr <caz<cr

26666664

37777775

Qcd1

Qcd2

Qcd3

Qcd4

Qcd5

8>>>>>><>>>>>>:

9>>>>>>=>>>>>>;

ð2Þ

Em~ Wm QmrzQmlzQmdrzQmdlð Þ

zX5

j~1

Njij QcnzQcdj

� �#ð3Þ

Fm~L Em

L xs~f0 xsð ÞW2

mzf1 xsð ÞWm

|X5

j~1

Nj ijzf2 xsð ÞX5

j~1

Nj ij

!2

ð4Þ

As pointed out by equation (4), the magnetic force can

be expressed as a sum of three basic terms. The first

one is exclusively attributable to the permanent

magnets; it depends on the armature position, and

its effect can be viewed as a diminution of the centring

spring stiffness. The second term is the most important

force contribution; it is linear with the coil currents,

and it is produced by the interaction between the

magnetic fluxes induced by the currents and the ones

induced by the permanent magnets. The third term is

exclusively attributable to the coils, but it tends to be

important only when high currents are considered [9].

3.2 Dynamics of the currents

If a partial short circuit is applied to the hth motor

coil, the dynamics of the currents is regulated by

equations (5) and (6). In particular, equation (5)

refers to all the active coils and to the section of the

failed coil that is still controlled by the servo-

amplifier, while equation (6) refers to the shorted

section of the failed coil. The resistances of the coils

are finally calculated taking into account the number

of their windings, equation (7)

Rc j ijzX5

k~1

Ljkdik

dtzKb _xxs~Ksa Rs icom j{ij

� �V j~1, . . . ,5 j=hz1ð Þ ð5Þ

Rc hz1ð Þihz1zX5

k~1

L hz1ð Þ kdik

dtzKb _xxs~0 ð6Þ

Rc j~Rc V j~1, . . . ,5 j=h; j=hz1ð ÞRc h~Rc 1{ Nsc

N

� �Rc hz1ð Þ~Rc

Nsc

N

8><>:

ð7Þ

3.3 Valve spool dynamics

Assuming that the centring spring is unloaded when

the spool is centred (i.e. xs 5 0), the spool displace-

ment dynamics is generally described by equation (8)

ms €xxszCs _xxszKs xs~FmzFfzFfr ð8Þ

where ms is the spool mass, Ks is the centring spring

stiffness, and Cs includes all the linear damping

effects (the viscous friction and the internal damping

of the centring spring), while Ff and Ffr respectively

represent the flow force and the sliding friction

force.

Concerning the flow force, it must be outlined that

the validation of the servovalve model (section 3.4)

has been obtained without flow regulation, so the

flow force has not been taken into account in the

servovalve model.

Regarding the sliding friction, an accurate model-

ling is necessary to obtain significant results from

the simulation of the flight control system built-in-

tests (section 4.1), since they are based on the

verification of the valve current-to-displacement

static response, which is heavily influenced by the

friction effects. Thus, the single-state elastoplastic

friction model [10] given by equations (9) to (11) has

been used

Ffr~{Kfr z ð9Þ

_zz~ _xxs{a _xxsj jz

zslð10Þ

a~

0 zj j¡zst

sin2 p zj j{zstð Þ2 zsl{zstð Þ

� �zstv zj jvzsl

1 zj j¢zsl

8>>><>>>:

ð11Þ

12 #





where z is the local deformation of the spool contact

area, Kfr is the contact stiffness, a is the stick-slip

factor, zsl is the local deformation during the slipping

motion (assumed to be constant), and zst is the

maximum local deformation during the sticking

motion (assumed to be zsl/2).

3.4 Experimental validation

The basic problem in the study of the short circuit

failure is attributable to the fact that the fault can

involve a limited part of the coil, and it is very complex

experimentally to reproduce this condition. This leads

to the necessity of modelling and simulation for

estimating the effects of the failure on the actuator

dynamics, but it also implies some kind of uncertainty

about the verification of the simulation results. On the

other hand, the experimental study of the total short

circuits is rather simple to realize, as the faults can be

reproduced without performing invasive tests on the

component. This aspect was important for the

research because parallel experimental activities were

carried out on the flight actuator, and others were

planned for the future.

For these reasons, the validation of the model has

been obtained only with reference to the total short

circuits, by testing an actual aircraft servovalve with

an electronic control unit specifically designed for

the artificial injection of the failures [11]. Figure 6

reports the working scheme of one of the four

electronic circuitries used for testing the fault-

tolerant servovalve. Three discrete switches have

been inserted: the closing of the switch A results in

injecting a total short circuit, while the closing of the

switches B and C results in engaging the servo-

amplifier with the coil.

Both the static and the dynamic performances of

the servovalve have been characterized, by testing the

component at different working conditions and with-

out flow regulation. Figure 7 shows the comparison

between the current-to-displacement curves obtained

using the hardware and the model at different working

conditions. Each curve refers to a test in which a

sinusoidal low-frequency command signal (¡imax at

0.02Hz) is provided to the direct-drive motor, while

the spool displacement and the coil current are

measured. It can be noted that the simulation results

satisfactorily match the experimental data, both in

normal condition and in the case of short circuits. The

model is able to reproduce the hysteresis amplitude,

which basically depends on the effects of the sliding

friction, as well as the overall shape of the current-to-

displacement curve, which is essentially attributable

to the non-linear relation between the magnetic force

and the armature position, equation (4).

Figure 8 shows the comparison between the hard-

ware and the model frequency responses of the

servovalve with reference to the normal condition

and in case of total short circuit failures. The results

are obtained by commanding the servovalve with 20

per cent of the maximum current and by defining the

output signal as a percentage of the spool full stroke.

The model generally matches the experimental

data, even if the phase errors tend to increase in the

high-frequency range if the failed conditions are

concerned. The short circuits lower the steady-state

asymptote of the frequency response (coherently

with the results shown in Fig. 7) and they also

reduce the servovalve bandwidth, which passes from

about 5 Hz to about 2.5 Hz when two short circuits

occur. The phenomenon depends on the fact that

the current generated in the shorted coils is opposite

to the ones circulating in the active coils, so that the

short circuits act as brakes in the spool dynamics.

Fig. 6 Scheme of the servovalve electronic control unit: (a) active coil and (b) shorted coil





The proposed model provided satisfactory results,

but further studies have been planned with the

purpose of obtaining a complete validation of the

modelling schemes in terms of reliability and the

robustness of results (e.g. Monte-Carlo statistical

analysis).

3.5 Numerical characteristics of the model

The servovalve model has been entirely developed in

the Matlab-Simulink environment, using the State-

flow charts for the reconfiguration of the model due

to failure events. The numerical solution demon-

strated to be stable and accurate using the Runge–

Kutta method (ODE 4) with an integration step of

18.5 ms. The need of this very small integration step

results from the simulation of event-driven dy-

namics such as the valve sliding friction (section

3.3) and the current dynamics in case of failures. In

particular, it must be outlined that, once defined

these simulation parameters, it is not possible to

simulate the servovalve dynamics introducing partial

short circuits with too-small or too-large extensions.

This limit depends on the fact that, for both cases,

Fig. 7 Model validation: servovalve current-to-displacement curves at different workingconditions





the magnetic inductance of one coil would be too

small, inducing too-fast current responses and, as a

consequence, numerical instability.

In the perspective of using the servovalve model

for real-time/hardware-in-the-loop applications, a

real-time executable code has been obtained by

means of the xPC Target tools, and the model results

to be suitable for the real-time execution on a

common off-the-shelf PC (AMD Athlon 2.1 GHz),

even if one PC must be completely dedicated to the

servovalve simulation.

4 FAULT DETECTION PROBLEMS ANDACTUATOR PERFORMANCE DEGRADATION

4.1 Simulation of the direct-drive servovalve on-ground built-in-test

In a fly-by-wire aircraft, the correct operational state

of each critical subsystem, such as a primary flight

actuator, must be assessed by the flight control

computers before and during the mission, through

check procedures known as built-in-tests. In parti-

cular, the basic objective of an on-ground built-in-

test is the detection of dormant failures, which

(depending on the failure type as well as on the

system availability requirement) can even lead to

aborted mission. On the other hand, the in-flight

built-in-tests aim to detect, isolate, and possibly

compensate the failures, to maintain the appropriate

levels of safety and performance [1].

The direct-drive servovalve of a primary flight

actuator is a critical component of a flight control

system, as the aircraft controllability directly depends

on the performances of its flight controls, which are

strongly influenced by the servovalve dynamics. For

this reason, specific on-ground built-in-tests are

carried out on the direct-drive servovalve, with the

purpose of detecting electrical malfunctions, mechan-

ical seizures, as well as the overall valve performance

(e.g. by testing the capability of driving the spool all

along its stroke with the minimum tolerable number

of active coils).

Concerning the electrical failures (section 2.2), the

identification of both the hardover and the open

circuit is quite reliable and simple to implement,

since a comparison between the demand current

and the actual one (two signals that are always

available in the control electronics) is sufficient. On

the other hand, a direct and reliable detection of a

short circuit could be obtained only by measuring

the current and the pin-to-pin voltage of a coil;

however, the latter signal is usually not available in

the common implementations [12].

Fig. 8 Model validation: servovalve open-loop frequency responses at different workingconditions





An alternative approach could be based on the

evaluation of the servovalve current-to-displacement

response related to the single coils (Fig. 7(b)), but

this procedure could imply the possibility of having

dormant failures owing to partial faults. This

problem has been pointed out by means of the

validated model of the servovalve, simulating an on-

ground built-in-test composed of the following

steps.

Step 1: all the coil currents are set to zero; after

that, a test current (iBIT) is sent to the coil 1 and a

static measurement of the spool position (xs1 BIT) is

recorded;

Step(s) 2 (3 and 4): step 1 is repeated commanding

the coil 2 (3 and 4);

Step 5: if the jth spool position measurement

(where j 5 1,…,4) is outside a fixed threshold (i.e.

|xs j BIT 2 xs nom| .Dxth), a malfunction is detected

in the jth coil; otherwise the servovalve passes the

check.

The failure coverage provided by this test clearly

depends on the position threshold (Dxth): in parti-

cular, the larger is the threshold, the bigger is the

undetected partial short circuit. Nevertheless, too-

small thresholds cannot be used, because the

current-to-displacement response of a direct-drive

motor depends on the temperature (the air gaps vary

with the thermal dilatation and the magnet proper-

ties are temperature dependent [8]), and the com-

ponent temperature during the built-in-test is not

predictable.

For example, once assumed iBIT 5 imax (xs nom <0.4xs max, Fig. 7(b)) and Dxth 5 0.1xs max, Figs 9 and 10

report the results of built-in-test simulations per-

formed with no failures and with increasing exten-

sions of a partial short circuit (involving 20 and 50

per cent of the total windings). Figure 9 shows the

spool displacement responses; it can be noted that, if

the fault involves 20 per cent of the coil, the test is

not able to detect the malfunction and a dormant

failure will result. It can be observed that this fault is

not severe and the performance degradation would

not be significant, but a small short circuit can

rapidly extend. Actually, Fig. 10 shows that the

current dynamics in the shorted section of the failed

coil is characterized by a transient that is as stronger

as the partial short circuit is smaller. Considering

that, when the motor works with all the coils, the

transients in shorted windings are much higher (the

peak value can reach up to 3–4 times the maximum

current), a rapid fault extension is expected after a

short period of operation.

For this reason, the performance degradation of

the flight actuator has been characterized with

reference to a completely extended failure (a total

short circuit) and, as the worst case condition for the

two-fail-operative servovalve, to the case of two total

short circuits.

4.2 Characterization of the actuator performancedegradation

The degradation of the flight actuator performances

owing to the short circuit failures has been char-

acterized through experiments.

As shown by Fig. 11, the flight actuator closed-

loop control is composed of two position feedbacks,

one acting on the servovalve spool (for enhancing

the component bandwidth) and the other on the

actuator rod. To obtain a basic characterization of

the system dynamics, simple proportional controls

have been used for both the loops.Fig. 9 Simulation of the servovalve built-in-test with a

partial short circuit: spool displacement

Fig. 10 Simulation of the servovalve built-in-test with apartial short circuit: current in the shorted coil





Figure 12 reports the closed-loop frequency re-

sponses of the flight actuator in different operative

conditions. The experiments – performed with a

command amplitude of about 1 per cent of the

actuator full stroke – highlight that in the normal

condition the actuator response is well damped and

exhibits a bandwidth (defined at 245u phase delay)

of about 5 Hz. If one short circuit is injected, the

variation of the actuator dynamics is negligible in

the low-frequency range (, 5 Hz), but it tends to be

more important at medium–high frequencies

(. 10 Hz), as the phase delay and the amplitude

attenuation are notably increased. If the actuator

works with two short circuits, the performance

degradation becomes significant, as the damping

becomes worse, the bandwidth lowers up to 3 Hz

and the phase delay is strongly increased (e.g. at

10 Hz, the two failures add about 40u phase delay to

the normal actuator response). These phenomena

can be justified and discussed with reference to

Fig. 13, where the closed-loop frequency responses

of the real-aircraft servovalve are compared with

those provided by the model. The results, which

generally provide a further validation of the model,

show the effects of the short circuit failures on the

closed-loop dynamics of the servovalve. In particular,

Fig. 11 Closed-loop control scheme of the flight actuator

Fig. 12 Degradation of the actuator frequency response owing to short circuit failures





the steady-state asymptote of the amplitude response

lowers of about 2 dB if two short circuits are injected,

which means that, in these conditions, the actuator

open-loop gain is reduced of about 20 per cent and, as

a consequence, the closed-loop bandwidth lowers.

Moreover, Fig. 13 points out that the servovalve

bandwidth (30 Hz in the normal condition) lowers

up to 10 Hz if two short circuits occur, so that the valve

dynamics strictly couples with that of the rod,

and additional phase delay results. Finally, it must

be mentioned that the frequency responses reported

in Fig. 13 have been performed on the isolated

servovalve with a command amplitude of 10 per cent

of the maximum valve opening, but, in the whole

actuator dynamics, the valve commands rapidly tend

to saturate if the servovalve response is not fast

enough; so additional servovalve phase delay must be

taken into account when the actuator works in the fail-

operative conditions.

4.3 Failure compensation

The characterization of the actuator performances

discussed in section 4.2 has been obtained assuming

that no control reconfiguration is applied after the

failures. On the other hand, once the effects of the

short circuits on the servovalve dynamics are known,

compensation techniques similar to the one pro-

posed by the authors in reference [13] or others can

be used, significantly reducing the actuator sensitiv-

ity to the failures. However, no compensation can be

applied if the failure is not detected, and this

problem is relevant to the case of short circuits.

Actually, during the in-flight built-in-tests, the

electrical health state of a servovalve is typically

judged [12] by comparing the current signal fed back

by the servo-amplifier with a reference value

provided by a servo-amplifier model (in-line mon-

itoring) or obtained by a comparison with the signals

coming from the other lines (crossline monitoring).

Unfortunately, these techniques are not effective for

detecting the short circuits. As shown in Fig. 3, the

servo-amplifier connected to the failed coil feeds

back a current as well as the one connected to the

active coil, and, as the commands are the same, the

current sensed in the failed line is almost identical to

the ones related to the active lines. For these

reasons, both the inline and the crossline monitoring

are not able to detect the short circuits, and no

compensation can be applied.

A possible solution can be obtained by an inline

monitoring based on an accurate and real-time

Fig. 13 Degradation of the servovalve closed-loop frequency response owing to short circuitfailures





executable model of the direct-drive servovalve,

which should be able to provide a precise correlation

between the current commands and the spool

position, allowing identification of the number (even

if not the location) of short circuits, and to imple-

ment the system compensation.

5 CONCLUSION

The dynamics of a fault-tolerant direct-drive servovalve

for primary fly-by-wire actuators has been studied –

developing and validating through experiments a high-

fidelity model that is able to simulate the component

behaviour in partial and total short circuits to the direct-

drive motor coils. The model has then been used for

simulating a typical on-ground built-in-test procedure,

obtaining a servovalve dormant failure related to a short

circuit involving the 20 per cent of the windings of a coil.

Simulation results also highlighted that the current

circulating in the shorted section of a coil can reach

values of about 3–4 times the maximum allowable

current, thus implying a rapid fault extension after a

short period of operation. For this reason, the experi-

mental characterization of the flight actuator perfor-

mance has been obtained in the case of one and (as the

worst case condition) two total short circuits. Experi-

ments highlighted that the short circuits almost halve

the actuator bandwidth and also cause a reduction of

the response damping. The phenomena have been

correlated with the effects of the failures on the isolated

servovalve dynamics that reduce both the actuator

open-loop gain and the servovalve bandwidth. Finally,

the problems related to the in-flight detection and

compensation of the short circuits with the traditional

monitoring techniques have been discussed, pointing

out the potential effectiveness of an inline monitoring

based on an accurate and real-time executable model of

the direct-drive servovalve dynamics.

REFERENCES

1 Pratt, R. W. Flight control systems: practical issuesin design and implementation, 2000 (Institution ofEngineering and Technology, Stevenage).

2 Miller, F. G. Direct drive control valves and theirapplications. Proceedings of the IMechE Interna-tional Conference on Aerospace hydraulics andsystems, September 1993, pp. 1–16.

3 Steed, D. J. Direct drive actuation for primary flightcontrol. Proceedings of the IMechE InternationalConference on Aerospace hydraulics and systems,September 1993, pp. 17–22.

4 Trutt, F. C., Santa Cruz, C., Kohler, J. L., andSottile, J. Prediction of electrical behavior indeteriorating induction motors. IEEE Trans. Ind.Applic., 1993, 29, 1239–1243.

5 Chen, H. and Zhang, C. Analysis of electrical faultsfor switched reluctance motor driver. In Proceed-ings of the 3rd World Congress on Intelligentcontrol and automation, 2000, Vol. 1, 666–669.

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7 Fitzgerald, A. E., Kingsley, C. Jr, and Kusko, A.Electric machinery: the processes devices and sys-tems of electromechanical energy conversion, 3rdedition, 1971 (McGraw-Hill, New York).

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9 Di Rito, G. Experimental validation of theoreticaland numerical models of a DDV linear force motor.Proceedings of the 3rd FPNI-PhD Symposium,2004, pp. 105–114.

10 Dupont, P., Hayward, V., Armstrong, B., andAltpeter, F. Single state elastoplastic frictionmodels. IEEE Trans. Autom. Control, 2002, 47(5),787–792.

11 Denti, E., Di Rito, G., and Galatolo, R. Real-timehardware-in-the-loop simulation of fly-by-wireflight control systems. 25th International Congressof the Aeronautical Sciences (ICAS), 2006.

12 Bonnice, W. F. and Baker, W. Intelligent faultdiagnosis and failure management of flight controlactuation systems, NASA Contractor Report 177481,1988.

13 Di Rito, G. and Galatolo, R. Modelling thedynamics of a direct-drive servovalve for fail-operative primary flight control actuators. Proceed-ings of the 3rd International Conference on Recentadvances in aerospace actuation systems andcomponents, 2004, pp. 67–72.

APPENDIX

Notation

Cs spool damping (kg/s)

Em field magnetic co-energy (J)

Ff flow force (N)

Ffr sliding friction force (N)

Fm magnetic force (N)

icom j current command to the jth coil (A)

ij current circulating in the jth coil (A)

j index related to the direct-drive

servovalve coil

Kb back-electromotive force (V s/m)

Kfr contact stiffness (kg/s2)





Ks centring spring stiffness (kg/s2)

Ksa servo-amplifier gain

Ljk magnetic inductance of the jth coil

related to the current circulation in

the kth coil (H)

ms spool mass (kg)

N total number of coil windings

Nsc number of shorted coil windings

Rc j resistance of the jth coil (Ohm)

Rs sense resistance (Ohm)

xs spool displacement (m)

z local deformation of the contact area

(m)

zsl local deformation during the slipping

motion (m)

zst maximum local deformation during

the sticking motion (m)

a stick-slip factor

Qcdj dispersed magnetic flux linked to the

jth coil (Wb)

Qcn net magnetic flux linked to armature

and coils (Wb)

Qmdl dispersed magnetic flux linked to the

left permanent magnet (Wb)

Qmdr dispersed magnetic flux linked to the

right permanent magnet (Wb)

Qml magnetic flux linked to the left

permanent magnet (Wb)

Qmr magnetic flux linked to the right

permanent magnet (Wb)

Wm magnetomotive force of the perma-

nent magnet (A)

Ral magnetic reluctance of the armature

axial gap (left side) (A/Wb)

Rar magnetic reluctance of the armature

axial gap (right side) (A/Wb)

Rca magnetic reluctance of the axial

dispersion path of a coil (A/Wb)

Rcr magnetic reluctance of the

radial dispersion path of a coil

(A/Wb)

Rmd magnetic reluctance of the disper-

sion path of the permanent magnet

(A/Wb)

Rmi internal magnetic reluctance of the

permanent magnet (A/Wb)

Rrc magnetic reluctance of the armature

radial clearance (A/Wb)



