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Vol.107 (3) September 2016 SOUTH AFRICAN INSTITUTE OF ELECTRICAL ENGINEERS 167

FAILURE ANALYSIS OF METAL OXIDE ARRESTERS UNDERHARMONIC DISTORTION

P. Bokoro∗ and I. Jandrell†

∗ University of Johannesburg, Dept. of Electrical Engineering Technology, 37 Nind Street, Universityof Johannesburg, South Africa E-mail: [email protected]† Faculty of Engineering and the Built Environment, Private Bag 3, Wits 2050, South Africa E-mail:[email protected]

Abstract: The probability of accelerated degradation or reduced time to failure of metal oxide arresters(MOA), when continuously exposed to distorted ac conduction, is analysed in this study. Metaloxide-based surge arresters of similar size and electrical characteristics are tested using accelerateddegradation at elevated temperature and distorted ac voltage stress. The three-parameter Weibullprobability method is applied to analyse the obtained time to failure distribution. The Fourier series’expansion is also relied upon in order to evaluate the content of the harmonic resistive components ofthe measured leakage current. The results obtained indicate that for 6.24% and 5.58% content of the 3rd

and the 5th harmonic component, respectively embedded in the applied voltage stress, the probability ofreduced time to failure or accelerated degradation is found to be 58.93 %, and the mean life reductionobtained is just above 40%. These results correlate with the pronounced shift of the U − I curve as wellas the increase in the magnitude of the respective harmonic resistive current components of the arrestersamples.

Key words: Time to failure, metal oxide arrester, voltage harmonics, resistive current, reliability.

1. INTRODUCTION

Metal Oxide arresters (MOA) experience electricaldegradation or ageing as a result of continuous ac ordc conduction [1-3]. The degradation phenomenonrefers to irreversible change in electrical and physicalproperties of the MOA components [4, 5], and consistsof one of the most encountered failure modes that affectmetal oxide-based arrester components [6, 7]. Withthe prevalence of harmonic-producing loads in modernelectrical networks, MOA devices are likely to becontinuously exposed to voltage and current harmonicfrequencies. Previous studies described in [8-11], haveindicated the impact of voltage harmonics on the leakagecurrent-based condition assessment of metal oxide-basedsurge arresters. The major shortcoming in these studies isthe probable aggravated degradation or failure process ofthese surge protection devices as a result of voltage andcurrent harmonics in the power system. In the presentwork, two commercially-sourced sets of MOA are testedand analysed. Each set consisted of 60 units of similarphysical and electrical characteristics. These arrestercomponents are subjected to accelerated degradation test at135oC for a time period of 96 hours. The applied voltagestress consisted of distorted waveform whose fundamentalcomponent amounted to 85 % of the ac rated breakdownvoltage (0.85U1mAac) of the MOA. The 3rd and 5th

harmonic voltage components are found to be prominentin the applied voltage stress, and their percentage contentis considerably changed when the harmonic source usedin this study is removed from the circuit. The resultsobtained show that arrester samples subjected to voltagestress, containing higher harmonic content, exhibit higherprobability of failure or degradation, and significant

rise in the magnitude of the resistive harmonic currentcomponents.

2. EXPERIMENTAL WORK

For the metal oxide-based arresters used in this study, twoseparate types of test set up are applied: the accelerateddegradation test at elevated voltage and temperature, whichemulates real life arrester deterioration process [12], andthe dc voltage-current (U − I) test for the purpose ofreference voltage (Un) measurement, which enables thedegradation or failure status of arresters to be verified.

2.1 The accelerated degradation test

This test regime essentially consisted of the followingcomponents: a heat chamber or mini-furnace, a 50 Hzac supply voltage, the triac - based voltage controller(harmonic source), a resistive load, high-temperatureconductors, data logger units and FN 2090 single-phasemulti-purpose filter. The heat chamber consisted of theNabertherm P330 with 9 settable heating programs orcourses (P1-P9), and 40 time-segments grouped in blocksof 10 (A-J)[13]. Each block is made up of 4 time-segments(2 ramp and 2 holding times). A heating program could beformed of one or more blocks. In this study, the desiredtemperature is reached at the first holding time-segment(t2B) of block B, which is assigned a time value of96 hours and a temperature of 135oC. The subsequenttime-segments are therefore assigned a zero value. Atthe end of the degradation time thus set, the unit willautomatically switch off and enter the waiting time orcooling mode. The heating program used in this study isshown in figure 1.

Vol.107 (3) September 2016SOUTH AFRICAN INSTITUTE OF ELECTRICAL ENGINEERS168

Figure 1: Heating Program

When the triac-based voltage controller is removed fromthe circuit, and the multi-purpose filter is connected: thenon-linear leakage current resulting from high conductionthrough MOA devices induced 1.89 % of the 3rd and2.5% of the 5th, which fall within the permissible levelof harmonic [14]. Upon the connection of the triac-basedvoltage controller, these harmonic components increasedto values beyond permissible level: 6.24 % and 5.58 %,respectively. The set up of the accelerated degradation testis provided in figure 2.

Figure 2: Accelerated Degradation Test Set up

Each test run accommodates five MOA samples connectedacross terminals, mounted in parallel on a concreteplatform inside the heat chamber. These terminals aresupplied through 1.5 mm2 high-temperature single-coresilicon cables (Silflex), capable to withstand 180oC [15].The supply voltage is controlled from an external timerunit set to trigger when the chamber reaches 135oCof temperature. To prevent any event of short circuit,protective fuses rated 250 V; 0.125 A are connected inseries with each test sample. A three channel K5020and 2 x MT250 data logger units are connected in sucha way that voltage events across each sample are recorded.The TDS 1001B two-channel Tektronix and the 4-channelRigol digital scopes are used to monitor the supply voltageand to record the leakage currents, with the aid of 5/1A current transformers and a current probe. The failure

or breakdown times measured in both test conditions areextrapolated, using the Arrhenius model, to time values (ti)corresponding to long-term operation of these devices atservice condition: 40oC at maximum condition operatingvoltage (MCOV). Prior to the estimation of the shapeand scale parameters of the obtained time to failuredistributions, such as described in the IEEE guide forstatistical analysis of insulation breakdown [16]. Theapplied voltage and leakage currents are measured duringthe test, in comma separated values (CSV) format usingthe Rigol DS1204B digital scope. The Fourier seriesexpansion technique is applied to determine the magnitudeof the harmonic resistive current components in measuredleakage current waveforms. The U − I characteristiccurves of the tested samples are also plotted. The voltagestress applied to arresters (without external harmonics) wasprovided through FN 2090 single-phase and multi-purposefilter, in a bid to reduce harmonics from the mains. A83.3 kΩ resistive load was connected across the filterand outside the heat chamber. The time-domain of thecontinuous voltage stress applied under this condition isindicated in figure 3.

Figure 3: Applied Voltage Stress (no external harmonic source)

The harmonic components in the voltage stress resultingfrom non-linear current conduction through arrestercomponents are shown in figure 4. To introduceharmonic distortion from an external harmonic source, thetriac-based voltage controller is brought in to effect theswitching of the resistive load, the applied voltage stresstherefore experiences much higher harmonic distortion.This is shown in figure 5. The resulting voltage harmoniccomponents in the applied stress, when an external sourceof harmonic is depicted in figure 6.

2.2 Description of Arrester Samples

The overvoltage protection samples used in this workconsisted of low-voltage MOA units with 20 mm ofdiameter size. The mean resistance (Rmean), inductance(Lmean) and capacitance (Cmean) of the varistor arrestersare measured at room temperature, using an ELC-131DLCR meter. The breakdown voltage (U1mAac) as wellas the MCOV (Uac) of the MOA devices are obtainedfrom the manufacturer [17]. The electrical properties orcharacteristics of the tested arresters, such as specified in


Figure 4: Frequency Components of the applied voltage stress(No external harmonic source)

Figure 5: Applied Voltage Stress (with external harmonicsource)

the manufacturer’s data book, are provided in table 1.

2.3 U − I Measurement of MOA Samples

The U − I characteristic curve of MOA samples ismeasured at room temperature using a variable dc source.The arrester is therefore connected across the output. Avolt and current meter are respectively connected acrossand in series with the device under test. The measurementpoints obtained are subsequently used to plot the U − Icharacteristic curve. This test is conducted before andafter accelerated degradation test at room temperature. Thereference voltage (Un) and the standby leakage current(ILo ), which is defined as the current measured at 85% ofthe reference voltage [18, 19], could therefore be obtained.The mean U − I curve of the samples before and afterdegradation test are provided in figure 7.

Figure 6: Frequency Components of the applied voltage stress(With external harmonic source)

Table 1: Electrical Characteristics of Arrester SamplesCharacteristics Values UnitsRmean 5.6 MΩLmean 18.04 HCmean 1225 pFU1mAac 205 VUac 130 V

3. WEIBULL PLOTS

3.1 Time data obtained:

The procedure followed to obtain the time to failuredistribution and the resulting probability functions issummarised in figure 8. The failure times (ti) measuredduring accelerated degradation tests and the change in thedc reference voltage (Un), are used to isolate the survivedand spoiled components from the failed arrester samples,on the basis of the following conditions:

1. Failed samples: ti ≤ t2B and Un ≥ 5%

2. Survived samples: ti = t2B and Un < 5%

3. Spoiled samples: ti < t2B and Un < 5%

The classification of arrester samples as result of the aboveconditions is shown in table 2. Using the Arrheniusaccelerating factor [20], the time-points measured forthe failed samples are extrapolated to unit valuescorresponding to equivalent operation of arresters atstandard temperature of 40oC.

[teq.40oC]i [hours] = ti ×2.5T2B−40

10 (1)

Where:


Figure 7: Mean U-I curve of Arrester Samples

Figure 8: Block Diagram of Failure Probability Analysis

[teq.40oC]i = time equivalent to arrester operation at 40oC orservice conditionT2B = the test temperature.

Each failure time is assigned a ranking number (i) whichcounts the number of failures in a given time, andultimately the number of failed arresters at any indicatedtime. This helps to determine the percentage cumulativefailure probability using White’s approximation [21, 22].

F(i,n)≈(

i−0.44n+0.25

)×100 (2)

Where:

F(i;n) = the percentage cumulative failure probabilityi = the number of failures in a given timen = the number of tested samples.

The Weibull cumulative distribution function (CDF) ofboth sample groups, can therefore be plotted on the basisof [teq.40oC]i and F(i,n). The CDF of arresters exposed to

Table 2: Classification of degraded SamplesSamples No harmonic

sourceWith harmonic source

Failed 27 37Survived 30 18Spoiled 3 5

no external harmonic source and that of those subjected toan external harmonic source are shown in figures 9 and 10,respectively.

Figure 9: CDF (Sample without external harmonics)

Figure 10: CDF (Sample with external harmonics)

3.2 Adequacy of the Distribution:

To test the adequacy or the goodness of fit of thedistributions obtained, both the F(i;n) and [teq.40oC]i areassigned respective logarithmic expressions xi and yi asindicated in [23, 24]. These values are determined fromthe following equations:

xi = ln[− ln

(1− F(i,n)

100

)](3)

Where:

xi = the logarithmic expression of the percentagecumulative failure probability

And:

yi = ln(teq.40oC)i (4)

Where:

yi = the logarithmic expression of the time equivalent toarrester operation at 40oC or service condition.


The xi and yi values are used to determine the correlationfactor, which should be equal or higher than the Weibullcritical coefficient value (γ) for a curve to be deemedfit or adequate [23]. The correlation factor is calculatedusing the correlation function equation [25], expressed asfollows:

γ(xi,yi) =∑(xi − x) · (yi − y)√

∑(xi − x)2 ·∑(yi − y)2(5)

Where:

γ(xi,yi) = the correlation functionx = the mean logarithmic value of the percentagecumulative failure probability ( ∑xi

r )y = the mean value of the time equivalent to arresteroperation at 40oC or service condition ( ∑yi

r )r = the number of failed arresters.

The correlation factors of the distributions obtained, usingequation 5 are: γ1 = 0.955857 and γ2 = 0.958317. Basedon the Weibull critical coefficient values provided in [23],both curves show good adequacy or fit to the Weibulldistribution.

3.3 Estimation of the Weibull Parameters:

The least-squares regression method is used to determinethe slope m(xi,yi) and the c-intercept functions of theplotted curves [23, 24]. These functions are in turn usedto estimate the shape and the scale parameters. The slopefunction is determined in equation (6), as follows:

m(xi,yi) =∑(xi − x) · (yi − y)

∑(xi − x)2 (6)

Where:

m(xi,yi) = the slope function of the Weibull distribution.

Therefore, the shape parameter β is expressed as:

β =1

m(xi,yi)=

∑(xi − x)2

∑(xi − x) · (yi − y)(7)

Where:

β = the shape parameter of the Weibull distribution.

The c-intercept function is calculated using equation 8,given below:

c = y−m(xi,yi) · x (8)

Where:

c = the c-intercept function.

The scale parameter is therefore obtained using theexponential of the c-intercept function. This yields thefollowing expression:

α = expc = exp [y−m(xi,yi) · x] (9)

Where:

α = the scale parameter of the Weibull distribution.

Applying equations 6,7,8 and 9 yield the scale and shapeparameters of the time to failure distribution with andwithout harmonics to be determined: β1 = 0.98 and α1= 4167.6 hours, and β2 = 1.093 and α2 = 2746.5 hours.For both distributions, the minimum time to failure or thelocation parameter γ = 100.6 hours, has the same value.

4. FAILURE ANALYSIS

Based on the three-parameter Weibull distributionsobtained respectively for each set of degraded arresters(with and without harmonics), the probability densityfunction (PDF) can be determined and subsequentlyanalysed. Equation 10 is therefore applied:

f (t,β,α,γ) =βα·(

t − γα

)β−1

exp(− t − γ

α

)β(10)

Where:

f (t,β,α,γ) = the probability density fucnctionβα ·

( t−γα)β−1

= the hasard or failure rate function

exp(− t−γ

α)β

= the reliability function.

The PDF, the failure rate and the reliability functionsof MOA samples degraded with and without harmonicsare noted as follows: f1 (t), h1 (t), R1 (t), f2 (t), h2 (t)and R2 (t), respectively. To determine whether or notelectrical failure of arrester units is indeed accelerated,as a result of external harmonic content in the appliedvoltage stress, the reliability, the failure rate and the PDFof these components under the testing conditions could beanalysed. Therefore, the following statements apply:

1. The probability of one population of tested arrestersto experience longer time to failure over the otheris verified on the basis of the following probabilitycondition [25]:

Pr [t2 ≥ t1] =∞∫

0

f1 (t) ·R2 (t)dt > 0.50. (11)

Where:


Pr [t2 ≥ t1] = the probability that the population ofarresters degraded with harmonics may experiencelonger time to failure.

2. The mean time to failure (MTTF) of the twopopulations of degraded arresters could also beanalysed in a bid to determine the probability ofhigher survival rate or reliability of one populationover the other [26].

∞∫

0

exp(− t − γ

α2

)β2

dt >∞∫

0

exp(− t − γ

α1

)β1

dt (12)

3. The plots of the failure rate functions as applied tothe degraded arrester populations must therefore besuch that: h2 (t) < h1 (t). This could be graphicallydemonstrated.

The above statements are therefore analysed to assess anyprobability of accelerated failure as a result of harmonicsin the applied voltage stress.

5. LEAKAGE CURRENT ANALYSIS

Long-term exposure of varistor arresters to continuousvoltage stress generally lead to failure of these devices.This is usually diagnosed in terms of increased magnitudeof harmonic resistive component, and most particularlythe third harmonic resistive current (THRC) componentof the leakage current [27-30]. In order to assess thecontributions of each voltage harmonic frequency, theleakage current of the degraded arrester populations ismeasured and captured in CSV format of the RigolDS1204B digital scope. The time-domain waveform of theleakage current for arresters degraded without harmonics isshown in figure 11.

Figure 11: Leakage current waveform (Arresters degradedwithout harmonics)

Similarly, the time-domain waveform of the leakagecurrent for arresters degraded with harmonics is indicatedin figure 12.

Figure 12: Leakage current (Arresters degraded with harmonics)

The current waveforms measured could be expressedin terms of frequency components using Fourier series’expansion, given the periodical behaviour of these currentfunctions. This implies that:

i(t) =ao

2+

∞

∑k=1

(ak cosωkt +bk sinωkt) (13)

Where:

i(t) = the leakage current function

ao = 2T

T∫0

i(t)dt

ak = 2T

T∫0

i(t) · (coskt)dt

bk = 2T

T∫0

i(t) · (sinkt)dt

T = the period of the function.

In order to evaluate the terms of equation (13), thetime-domain curve of the leakage current is divided upinto 20 equal time-intervals between 0 and T . Thecurrent values corresponding to the time points are sourcedfrom the CSV measurement of the leakage current. TheFourier expansion of the leakage current function definedin equation 13 could be rewritten as follows:

i(t) =Io

2+

∞

∑k=1

√2Iksin(ωkt +φk) (14)

Where:

Ik =√

a2k +b2

k = the RMS value of i(t)

ωk =2kΠ

T = the angular frequencyφk = arctan ak

bk= the respective phase angle for the kth

current harmonic frequency component.

The Fourier expansion such as defined above is alsoextended to the applied voltages measured across MOA


samples during degradation test. This yields the followingvoltage expression:

v(t) =Vo

2+

∞

∑k=1

√2Vksin(ωkt +θk) (15)

Where:

Vk =√

a2k +b2

k = the RMS value of v(t)

ωk =2kΠ

T = the angular frequencyθk = arctan ak

bk= the respective phase angle for the kth

voltage harmonic frequency component.

Since the power losses resulting from distorted voltageacross MOA units could be estimated by the sum powerlosses of each components [31]. The instantaneouspower p(t) could therefore be expressed as the product ofequations (14) and (15). Disregarding the dc componentsin both (14) and (15), the following equation is obtained:

p(t) =∞

∑k=1

√2Iksin(ωkt +φk) ·

√2Vksin(ωkt +θk) (16)

Where:

p(t) = the instantanous power.

Developing further equation (16) yields the active orresistive component (PR) of the total power absorbed bythe arrester:

PR =∞

∑k=1

VkIkcos(θk −φk) (17)

Where:

PR = the average power.

The magnitude of the fundamental and harmonic resistivecomponents constituting the leakage currents could beeffectively determined using the expression Ikcos(θk −φk).The fundamental, the third and fifth harmonic resistivecurrent components are shown in figure 13.

6. RESULTS AND DISCUSSION

Based on the mean U − I characteristic curve obtained,arrester population degraded with harmonics proved tohave the lowest decrease (100V ) in the reference voltage(Un) measured at 1 mA dc and at room temperature.Whereas arresters degraded without external source ofharmonics have experienced a decrease of about 150 Vin the reference voltage. This implies that the electricalstability of arresters degraded with harmonics is severelycompromised as shown in figure 7. Based on the

Figure 13: Resistive Current Components before and afterharmonics

conditions of failure stated above, it could be observed that61.67 % of the MOA population degraded with externalharmonics in the voltage stress experienced breakdown,as compared to 45% when the voltage stress contains noexternal harmonic. This could further be observed interms of the hasard or failure rate functions obtained. Thefailure rate function of arresters degraded without externalharmonics h1 (t) is observed to be sharply increasing from0 to 2.46 × 10−4 failures per hour in the time intervalt ∈ [100.6;500], before decreasing from 2.46 × 10−4 to2.35× 10−4 failures per hour across the time interval t ∈[500;4500]. For arrester samples degraded with externalharmonics, the failure rate function h2 (t) is observed tobe sharply increasing from 0 to 3.33× 10−4 failures perhour across the time interval t ∈ [100.6;500]. Across thetime interval t ∈ [500;4500], the failure rate h2 (t) indicateda lower rate increase from 3.33 × 10−4 to 4.16 × 10−4

failures per hour. The decreasing failure rate impliesearly failure of arrester components in the life cycleof these components under standard service condition,while the increasing failure rate suggests a wear out ofthese protective devices. This therefore indicates that forthe time-interval [100.6,4500], the failure rate arrestersdegraded without external harmonics is consistently lowerthan that of samples subjected to harmonics. This impliesthat the statement: h2 (t) < h1 (t), for t = [100.6;4500] isnot true and the opposite statement: h1 (t) < h2 (t), for t= [100.6;4500] is therefore true. The failure rate graphsh1(t) and h2(t) are shown in figure 14.

Figure 14: Failure rate function before and after harmonics

An analysis of the reliability function or the survival prob-ability graphs obtained across [100.6;4500] time-interval,for both observed populations shows that R1 (t) decaysfrom 100 % to 36 %, while R2 (t) changes from 100 %


to 17 %. Moreover, for each point in time belongingto interval [100.6;4500], it is observed that R1 (t) isconsistently less than R2 (t). Therefore, the lowerreliability or probability of survival for arresters degradedwith external harmonics, as compared to those degradedwith no external harmonics can be associated to the contentof harmonic distortion in the voltage stress. This impliesthe relationship: R1 (t) > R2 (t), for t = [100.6;4500].The reliability functions of both populations are plotted infigure 15.

Figure 15: Reliability functions

The probability of one population of tested arresters toexperience longer time to failure over the other, such asdetermined in equation 11, yields the probability value(Pr = 0.4107 or 41.07%), which is obviously less than0.50 or 50%. This suggests that the time to failure or theprobability of arresters, subjected to external harmonics,experiencing longer time to failure than those degradedwith harmonics is 41.07%. Since the probability sum:Pr [t2 ≥ t1] + Pr [t1 ≥ t2] = 1. This therefore impliesPr [t1 ≥ t2] = 1 - Pr [t2 ≥ t1] = 0.5893 or 58.93%. Thissuggests that the stated relationship: P [t2 ≥ t1] > 0.50cannot be true, and consequently the probability ofarresters, degraded without external harmonics, to survivelonger is actually 58.93% higher.The MTTF obtained for arresters degraded withoutharmonics is found to be 4252.65 hours as opposed to2512.81 hours for those degraded with external harmonics.This demonstrates that metal oxide arresters degraded withexternal harmonics experienced 40.91% reduction of theirlifetime, and therefore demonstrated lower reliability.The PDF curves obtained in figure 16 show higher densityof failure for the arrester components subjected to externalharmonics. The magnitude of the fundamental, the 3rd and5th harmonic resistive current components obtained are:0.013 mA, 0.005 mA, 0.003 mA for arresters degradedwithout harmonics and 0.137 mA, 0.104 mA, 0.082 mAfor those degraded with external harmonics, respectively.The resistive current component before harmonic injectionis therefore 0.0143 mA and 0.191 mA after injectionof external harmonics. An increase of 92.51% in theresistive current of the arrester samples degraded in thepresence of external harmonics was therefore observed.

Figure 16: PDF of degraded arrester populations

It could also be noted that the fundamental and theTHRC make up at least 90% of the total resistive currentmeasured before and after harmonics injection at thesame operating temperature. This shows correlationbetween the increase in the resistive current and thehigher probability of failure, the reduced MTTF, thehigher failure rate and lower reliability as well as thesevere shift in the U − I characteristic curve, in arresterpopulations subjected to external harmonics. The increasein the resistive current directly translates into increasedpower losses being absorbed by arrester components,which will therefore quicken the thermal runway process.Since the test temperature which represents the device’soperating environment was kept constant, the rise in theresistive current and the subsequent high power lossesexperienced, when external harmonics are injected, couldtherefore be attributed to the influence of harmonic voltagefrequencies on the overall continuous biasing effect ofthe applied voltage stress. Furthermore, the study onthe degradation mechanism of metal oxide-based arrestersdescribed in [4], revealed that the degradation of thesearresters consists of the resultant effect of the breakdownof several millions of individual grain boundaries atthe following voltage levels: 3.02 V and 3.11 V pergrain boundary, for monotonous and non-monotonousageing process respectively. This therefore suggests thatthe reduced time to failure and the increased resistivecurrent, observed in MOA samples subjected to externalharmonics injection, results from the contributing effect ofvoltage harmonic components on the breakdown voltagebetween individual grain boundaries. This demonstratesthat voltage harmonics could be regarded as aggravatingfactors of the long-term degradation phenomenon of theseovervoltage protective devices. The third harmonic voltagewill, by virtue of its magnitude, be the second voltagecontributor to the MOA microstructure disintegration,hence to the accelerated degradation.

7. CONCLUSION

For the purpose of this work, similar arrester units aresubjected to accelerated degradation test, at elevatedtemperature and voltage. The voltage applied is embeddedwith harmonics. The observed time to failure or lifeexpectancy represents the behaviour pattern of these


overvoltage units, when operating at normal servicecondition with distorted voltage for a long period of time.The resistive current components as well as the U − Icharacteristic curve confirm the the degradation status ofarrester units. The following findings are obtained:

1. Continuous exposure of metal oxide-arresters todistorted ac voltage is prone to aggravate thedegradation or failure process of these surgeprotection units.

2. The reduced time to failure or lifetime of MOA units,continuously operated under distorted ac voltage,could be attributed to additional watt loss experiencedin these devices which result from increased harmonicresistive current conduction.

3. Harmonic components embedded in the voltage stresscontribute to the biasing effect of MOA units.

These findings imply that the presence of harmoniccomponents in the voltage across arrester units willfast track the disintegration process of the intergranularboundaries of MOA arresters. This therefore explainsthe higher probability of electrical failure or reduced lifeexpectancy of metal oxide-based arresters under harmonicdistortion conditions. The develpment of new oxideadditives, capable to decelerate intergranular disintegrationunder the effect of voltage stress, could be recommendedas one of the directions for future designs of MOAs.

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