aech chapter 6

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Section II Bare Aluminum Wire and Cable

Chapter 6

Operating Performance and Problems

Operating problems occurring in installations of bare

overhead conductors are of several kinds. Only those related to the conductors themselves are considered herein.Such matters as voltage drop', system regulation, transients and calculation of probable short-circuit currents arein the province of the system electrical engineer and be

yond the scope of this book. Subjects covered' in thischapter include the ability of the conductor to withstandshort circuits and their related mechanical forces, the extent that emergency overloads may be carried withoutserious damage and the effects of arcing-burndown. Reference is also made to aeolian vibration and conductorgalloping with a brief description of devices that reducetheir effects.

The terms used herein relating to overload matters areas follows:

Thermal Limit (as associated with steady-state overloadconditions): The maximum temperature at which a con

ductor can operate continuously yet maintain the minimum tensile properties established by the manufactureror the user.

Arc-Current Burndown: Rapid failure caused by the heatof an arc on the surface of the conductor, accompaniedby the heat effect of current.

Fault-Current Burndown: Failure caused by overheatingas a result of a current overload. The conductor strengthdecreases sufficiently to cause tension failure.

Fault-Current Limit: The current (temperature) andtime combination which produces the maximum acceptable loss in conductor mechanical strength.

Current Values: Unless otherwise stated, all current valuesused in the discussion of overload conditions are in terms

of rms symmetrical amperes.

>I< Applying to bare transmission and distribution circuits only. Thecritical voltage-drop limitations of the National Electrical Code relatingto circuits under NECjurisdiction are mentioned in Sections 21O-19(a)and 21S-2(b) of the NEe, and the methods of computing drop orobtaining it from industry-supplied tables are described, applying toconductor sizes used mostly for interior circuits.

Short-Circuit Performance

The ampacity data in Chapter 3, Figs. 3-11 to 3-15,apply to steady-state normal operation for bare ACSRand all-aluminum conductors for temperatures up to1000C (600C rise over 400C ambient). This temperature

is frequently used for 1350-H19 conductors since thestrands retain approximately 90 percent of rated strengthafter 10,000 hours at temperature. (See Fig. 6-3.) ForACSR the strength is even less affected because the steelcore is essentially unaffected at these temperatures.

Short circuits in a power system can result in extremelylarge currents in conductors from the time of fault initiation until its interruption by the protective device, such

as circuit breaker or fuse. With modern relaying, the duration of the 60 Hz fault current is usually only from 3 to20 cycles for transmission circuits but may be longerfor distribution lines. I f the circuit is immediately reestablished by automatic reclosure and the fault has not

cleared, the total fault-current time will be the sum ofthe interrupting times.

Heating will generally be more rapid than cooling, andloss-of-strength estimates would require integration of thetemperature-time curve for temperatures above the arbitrary "damage" level. However, as temperature is notmeasured, a useful and practical alternative is to use the

current-time product and neglect the temperature slopes.When limits have been established, the time in which thefault must be cleared can then be determined.

In establishing suitable fault-current limits, 340'C hasbeen selected as the maximum temperature for all-aluminum conductors since momentary exposure to this tempera

ture does not result in a significant loss of strength. ForACSR or A WAC conductors with sizeable steel content(not the 18/1 or 36/1 strandings) an upper limit of645'C represents the threshold of melting for aluminumwith the sleel expected to supply the needed mechanicalstrength. The curves of Figs. 6-1 A, B, and C apply thiscriteria using an average specific heat and assume noheat loss from the aluminum strands during the shortduration of the fault current. Figs. 6-2 A, B, and C dothe same for ACSR conductors.

6-1



bare aluminum wire and cable

Adjustments for 6201-T81and ACAR Conductors

Values from Fig. 6-1 may be adapted to 6201-T81 andACAR conductors by applying suitable multiplying factors. Usually the value that is specified as the estimatedfault current is the known quantity, and the correspondingtime is found that wiU cause the upper temperature limitto reach 340°C over 40'C ambient for 61.2 percent lACS

conductor, thereby enabling the current-limiting devicesto be properly set. For other conductors, the time forthe 1350-H 19 conductor is multiplied by factors as below:

For 6201-T81 conductor, multiply by 0.903For ACAR conductor, see the applicable portion of

the following example:

Examples: Assume 500 kernil conductor and 20,000 rIDS 60 Hz faultcurrent. As this conductor size is not sh.own by Fig. 6 ~ 1 , the time isobtained by interpolating between values for 417 kcmil and 566.5 kcmHto 2.80 sec for 1350 H19. Then for 6 2 0 1 ~ T 8 1 it will be 2.80 x 0.903, or2 53 sec.

Fo r 241 13 ACAR, the time wiJt be

(2.80 X 0.65) + (2.51 X Q.J5) 2.71 sec.

Adjustmentjor Upper Temperature Limit

Whereas the upper-limit temperatures specified in Figs.6·1 and 6,2 are suitable for bare overhead conductors,there are conditions where a lower temperature, limit is

advisable, such as when the bare cable is confined inswitchgear or in switching compartments. Other con,dilion" such as the use of soldered, copper terminal pads;also may warrant a lower temperature limit. Multiplyingfactors for these conditions are as follows:

Multiply time from Fig. 6-1 by

For 1350-H19 6201,T81Upper Limit

3000C 0.903 0.814250'C 0.771 0.691200'C 0.621 0.559

and multiply time from Fig. 6-2 by

For ACSRUpper Limit

500'C 0.845400"C 0.721300°C 0.556

For 6201-T81 and ACAR, apply these factors afterapplying those as listed in the preceding section.

Arcing

Caution must be exercised in applying the fault-currenttimes, as described, for relay settings of protective deviceson distribution lines that may be subject to arcing burodovvn. Arcing locally cuts into the conductor quickly in

such cases. For example, a /1;0. 4/0 AWG 6/1 ACSRunder 1700·1b tension has arcing burndown time of 10 t14 cycles (.167 to .233 sec.) at 15,250 amp, whereas thjault·current limit time (there being no local arcing)1.6 sec for that current, under assumptions applying tFig. 6,2. Also see Table 6,1. For the usual transmissioline, or those at the higher distribution voltages, relaco,ordination on the basis of fault· current limit tim

usually is satisfactory, but for lower distribution voltagein metropolitan environments consideration should bgiven to arcing burndown.

Table 6-1 contains representative data from arcing tesconducted with the conductor under tension.

While arcing failure times are so short that little if anchange in tension can occur prior to failure, high faucurrents can heat the entire line. The reSUlting increasin sag can establish contact with ground or other condutor, initiating an arcing problem. Clearances can, therefore, be as sigoificant a constraint on maximum acceptabcurrent as is conductor strength.

Arcing Effects

Aluminum conductors resist damage by arcing bettethan conductors of other metals because the arc tends tcause less pitting and surface metal melting. Whesubjected to arc currents, the aluminum conductor sUlfacfrequently shows only a removal of sheen, slight roughening, and change of color over a considerable area. Theffect described applies to arcs of less intensity than thosthat produce arc-current burndown. However, the advantage of aluminum in this respect aids measurably ireducing operating costs, particularly for the smaller sizeof ACSR, in the many instances where small arcs resufrom flashovers, lightning, momentary contact with a trelimb, and the like.

From one group of tests, aluminum's resistancesurface damage from such minor arcing was evident wiarcs ranging up to about 78 cycles duration.'

Loss of Strength

The loss of conductor strength due to time at temperture is a cumulative effect. Heating due to short circuoccurrence should therefore be added to heating due tother circumstances to estimate the condition of the conductor. In actual practice, however, the total time of faucurrents is usually very small relative to emergency opeating time and is therefore igoored as an effeet on conductor strength. The temperature-time strength loss relationship is covered in more detail in the section oemergency loading (Chapter 12).

.. C. A. Martens, "Power Arc-Over On Overhead DistributioLines and New Developed Equipment for Protection AgainConductor Bumdown From That Cause," AlEE Technical PapNo. 4145, 1941.

6-2



operating performance and problems

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1\~ . I 2 / 0 A W G I

\ / I/O AWGJ~ ~ li.\ \/ 4.<l 2 AWG L~ 4.1 4 Awtl

1\1\

1'0\0 6AWG ILil Y.

BARE STRANDED ALUMINUM CONDUCTOR L ~ ~I\ : A ,\/

/\

\ \( O.0671m ) 2 ; t seconds. \ \ \ = =- - I m = area, cmils ~ r\ = amps, rms

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\ 1\

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1\

2

987

6

2 3456789 2 3456789 2 74 5 6 8 9

100 1,000 10,000 100,000

CURRENT IN AMPERES

Fig. 6-1A. Maximum fault-current operating limit for

stranded aluminum conductor. Upper temperature limit340°C, ambient temperature 40°C.

3

2

10 9 8 7

6

5

4

3

2

1.0 9 8 7

en6

0

5 z04 '- 'L&J

3 en

z2

L&J

:IE

I-0.1

9 8 7

6

5

4

3

0.01

0.005

Note 1. Time plotted is that required fo r a given rms

fault current to cause conductor damage due to annealing.2. Graphs asSUme there is no heat loss in the conductor.

The curve for all aluminum conductors may be applied to alloy 6201-T81 and ACAR conductors by computing theequivalent 1350-H19 cross section. The current may then be determined by extrapolating for the computed crosssection USing Figs. 6-1, A, B, and C.

6-3




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, \ \954 000 em~ ~ I /195.000 em'l t

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~ ~ III 636.000 emil I

605.000 emil I

IL556.000 emil I/

1\ \1\ / / /411.000 emil

t\ \ 1 / '/' ~ / / / 391,500 emil

i\/ / /i\ I\r'\ K>(/V

336.400 emil t

265.800 em , L,

\ l\ 1\\ ~ V4/0 AWG I

r\ 1\ 1\ 1\ 3!Q AW I

",

\ 1\', \ 1

!x \/.\\' \

BARE STRANDED ALUMINUM CONDUCTOR "- 1\ X \ \ \ \ , \\ \ \ ~ \ \ j \

J.-

1,000

It =

2

I I I 11111( O.067Jm ) 2 ; t seconds

I m = area, cmils

I = amps, rms

10,000

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j\\\ ' ~ ~ ~ ' 1\

t\"0 ~ ~ \ I

~ 1\2 3 4 6 789

1,000,000

CURRENT IN AMPERES

Fig, 6-1B. Maximum fault-current operating limit for Note: 1. Time plolted is that required for a given rmstranded aluminum conductor, Upper temperature limit fault current to cause conductor damage due to annealin340°C, ambienltemperature 40°C. 2. Graphs assume there is no heat loss in the conducto

The curve for all aluminum conductors may be applied to alloy 6201-1'81 and ACAR conductors by computing thequivalent 1350-H19 cross seclion. The current may then be determined by extrapolating for the computed crossection using Figs. 6-1, A, B, and C.

6-4



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- I = amps, rms

.

:

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1:590,000 em,r I

1/1,510,500 em,

,1,431.000 emi; I

IV1,351.000 emi I

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/11,272.000 emi

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1/i,}92,500 emir

1/(,,vI I

LlI),OOll em,r I1/

1.033.SOIl cmll

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5 6 I 4 5 9 1 4

100,000

CURRENT INFig. 6-IC. Maximum fault-current operating limit forstranded aluminum conductor. Upper temperatllre limit340°C, ambient temperature 40°C.

3

2

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6 5

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1.09 8 7

6 '":>5 z

04

<>.....

3 '"z

2

.....::IE

I-0.1

9 8 7

6

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1

2

0.01 9 8 1

6

AMPERES

Note: 1. Time plotted is that required for a given rmsfault current to cause conductor damage due to annealing.2. Graphs assume there is no heat loss in the conductor.

The Curve for all aluminum conductors may be applied to alloy 620J-T81 and ACAR conductors by computing theequivalent 1350-HJ9 cross section. The current may then be determined by extrapolating for the computed crosssection using Figs. 6-1, A, B, and C,

6-5




(I)

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21 - - - > -

,"

o . o o ~ 2 3456189 23456789 2 3456789

.00 ItOOO 10,000 100,000

CUR R EN T IN AMP ERE S

Fig. 6-2A. Maximum Fault-Current Operating Limit for Note: 1. Time plotted is that required fo r a given rmBare Stranded ACSR conductor. Upper temperature limit fault current to bring aluminum strands to the threshOl645 0C, ambient temperature4QoC. o/melting.

2. Graphs assume there is no heat loss in the conductor

6·6




!IQOO

CURRENT

Fig. 6·2B. Maximum Fault·Current Operating Limit forBare Stranded ACSR conductor. Upper temperature limit645 0 C. a m ~ i e n t temperature 400C.

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BARE STRANDED ACSR CONDUCTOR\

I = (O.0862m ) 2 'I = seconds

I m = area, cmils

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I 954,000 emil

I I79! ,000 emil.

II .500 emil

I 631.000 emil

1605.000 emil

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I 47: ,000 emil

13 97 .500 emil

I 336.400 emil

2&•.800 emil

II 4/0 AWG II

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\\

~ \ 1\~ ~ 1\

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IN AMPERES

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0005

00,000

Note: 1. Time plotted is that required for a given rmsfault current to oring aluminum strands to the thresholdofmelting.2. Graphs assume there is no heat loss in the conductor.

6·7




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BARE STRANDED ACSR CONDUCTOR

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V I1.510.500 emil

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~ \ V :.35,.000 cmil

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2 3 4

CURRENT IN AMPERES

Fig. 6-2C. Maximum Fault-Current Operating Limit fo r Note: I. Time plotted is that required fo r a given rmBare Stranded ACSR conductor. Upper temperature limit fault current to bring aluminum strands to the thresho6450C, ambient temperature 400C. oJmelting.

2. Graphs assume there is no heat loss in the conducto

6·8




where:

F = Pounds per linear foot of conductor

G = Multiplying factor, as in Table 6-2

I, and I, = Short-circuit current in each conductora-c symmetrical rms amp, or in d-c amp

d = Spacing between centerlines of conductors in

inches

Example: Assume a flat 3-phase circuit of 210 AWG-6/1 ACSRon 7-ft spacing. subjected to a fault current of 20,000 amp rms

symmetrical (line CT of Fig. 6-5). What is the average lateral forceexerted on the center conductor caused by an rms symmetrical

fully inductive fault current in the first offset loop (line OR of Fig. 6-5)

without allowing for mechanical damping, caused by inertia, elasticity.

and side-sway friction?

From table 6-2(d) the applicable multiplying factor G is 4.17.Applying Eq. 6-1, the average force F during the first current loop,assuming zero power factor is

4.17 X 5.4 X 20,0002

F=---------- 10.7 Ib per I t

7 X 12 X !O,

Under fault conditions, the mechanical action ofstranded conductors, which usually have very long spandistances compared to separation distances, is differentfrom the action of more rigid bus conductors describedin Chapter 13. The conductors can slap together violently-especially the subconductors of bundled conductorlines-and traveling waves move longitudinally along theline. Experience and testing have shown that this action isnot damaging to the mechanical strength of conductors or

insulators, but it must be carefully considered in the design and selection of spacers and dampers.

Emergency Loading

Transmission and distribution conductors are occasionally subjected to current overloads, due to emergency conditions, which produce temperatures beyond thenormal thermal limit. Coincidence of peak loads with highsummer ambients, shifting of additional loads to an already

loaded conductor, and use of high loadings to preventicing are some reasons for such overloads.

The question of what maximum conductor temperatures should be permitted for emergency operation depends on how much loss of strength is allowable and howlong the emergency-load temperature continues. The effectof heating is cumulative. As an example, if a conductoris heated under emergency loading for ten hours each year

for a period of ten years, the total effect is nearly thesame as heating the conductor continuously at that temperature for 100 hours.

Fig. 6-3 delineates the effect of time on 1350-H19 aluminum strand strength at three temperatures which areof interest to power engineers. The curves permit estimatesof the change in strength of conductors which have carried

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I I

L I

II /

/ /

II I

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100 90 80 70PERCENT REMAINING OF INITIAL STRENGTH

10,000

5,000

1,000

500

100 en

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50 1...=

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10 ...7 5

3

2

1

0.5

0.1

60

emergency overloads.Fig. 6-3. Time-temperature percent strength remaining in

An example is cited, based on the following assumptions IJ50-HI9 wire. Tensile tests made at room temperaturewhich should not be considered typical. after wire exposure to the indicated temperatures.

6-10



9000

8000

7000

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'"...5000'"

'";:;.."co"w

u;

Z 4000W I -

3000

2000

o 100 200 300

TEMPERATURE DEGREE C

Fig. 6-4. Reduction of breaking strength of aluminumand aluminum alloy stranded conductors of equivalentconductance. Breaking strength tests were made at roomtemperature after Vz hour exposure to elevated tempera-tures.

:

I000

I

I

.

!!

\i 246.9 kcmil6201-T81

I:

,

:

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,

~ i \\1--0

""AWG"\:211.6 kcmi!)

1350-H19

!

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!

i

:

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(1) Emergency conditions exist for 24 hours each year.

(2) The uselullife 01 the conductor is 30 years.


( 3) Maximum temperature for emergency condition,ISOoC (302 0 F)

(4) Conductor: 795 kcmil-37 str. 13S0-H19

At the end of 30 years, the conductor will have beenheated to 1500 C for 720 hours. Using Fig. 6-3 as a guidefor the estimate. the strength of 13.900 Ib would be reduced to approximately 10,600 Ib-about a 24"1. loss.

If the conductor were of the same size. but 2617 ACSR.the strength would be reduced from 31.500 Ib to 28.200 lb.or a 10 percent reduction. The advantage for ACSR isdue to the steel core, which is essentially unaffected by thetemperature range considered for emergency overloads.

Short time exposure to even higher temperatures canoccur, and Fig. 6-4 shows the effect of 112 hour of heating on similar conductors of three different aluminumalloys. Strength loss is rapid at temperatures above ISOoC.For momentary exposure to elevated temperature. thereis much less reduction in strength. The cumulative effectof a succession of short-time fault-currents during shortcircuits where high temperatures are possible plus emer

gency operation at lower temperature can cause conductorstrength loss which is of concern. However, knowledgeof the actual condit ions-current, time. ambient tempera·ture, wind velocity, conductor emissivity and the resultingactual conductor temperatures is seldom very precise. The"damage curves," Figs. 6-3 and 6-4, are also drawn fromdata having inherent variability. They therefore may beused only as a basis for a very approximate estimate ofthe actual condition of the conductor.

The creep rates at l50'C of the all-aluminum andaluminum alloy conductor are considerably higher thanthose of corresponding sizes of ACSR at the same tempera

ture. As was noted in Chapter 5, the creep rate used for

predicting 10-year final sags and tensions is based on thecreep rate at 60°F.

The analysis of the interaction of the thermal expansionrates, component stress levels and differing creep rates atelevated temperatures to determine the effect of high temperatures on final sags is very complex. High temperaturesfor time periods whleh may seem short in terms of thelife 01 the conductor can result in significant changes insag--especially for the conductor constructions which donot have significant proportions of steel. A method ofpractical calculations is presented in IEEE Paper TP

69-674-PWR by J. R. Harvey and R. E. Larson.A t}'Pical practice is to limit emergency load tempera

tures to a maximum of 125'C.

Vibration and Fatigue of Overhead Conductors'

An unprotected or improperly protected overhead conductor may undergo wind-induced vibrations under certain conditions to such an extent that fatigue failures ofstrands will develop at points of restraint or support.Similar failures have been observed at or near splices and

,. EPRI Handbook, "'Wind Induced Conductor Motion." contains anexcellent treatment of this subject.

6·11




TABLE 6-2

Multiplying Factors for Maximum Short Circuit Lateral Force Acting Upon Suspended Parallel Wires and Cables in various Arrang

ments Assuming Balanced Loading, in Terms of Direct, or of Symmetrical RMS Alternating Fault Current, Amp (Line C

Fig. 6-51 = I,m,

Arrangement of

circuit

Type ot circuit

and designationof location on

current·wave of

fault-producing

current

Conductor

upon which

force is

applied

Multiplying

factor G

(al

(bl

A

0

A

0

d

d

B

0

B

0

Direct current*

'-phase a-c

symmetrical

1-phase a-<:

asymmetrical

Aor B

Aor B

Aor B

1_0

2.0

8.0

(cl

(dl

A ~0'< /A ', Y

i B'-0--- -0l - d ~ 1

l_d_l_d_1

A B C

0 0 0

I

1·phase a·c

rms of first loop

3-phase a-c

asymmetrical

3-phase a-<:

rms of first loop

3-phase a-c

asymmetrical

same

3-phase a-<:

rms of first loop

same

Aor B

A, B, or C

A, 8, or C

B

Ao r C

B

Aor C

5.55

6.93

4.17

6.93

6.45

4.17

3.89

• Although steady·state direct-current implies that a multiplying factor of 1.0 is satisfactory, the transient and over·shoot at fau

initiation renders it common practice to use a factor of 2.0.

NOTES: All values assume a fully offset current wave in a fault of zero power factor without damping, or resonance effects fro

support vibration.

See NEMA BU-' for adjustment factors if fault-current power factor differs from zero, as determined by XlR ratio.

This arrangement of factors differs from that of ANSI (37.32 because it is usual practice to designate fault currents

apparatus and lines in terms of rms symmetrical amperes (I,m').

6-12




A

1\ ~Rr-

'- ,0

DQ.

"- ~ B B, - i :... - I- __~ " - - f-. r-- I = : : : ' ~ ' -===f> _-1-' -T__

I-J -- ,- -- f" - T

,,--/ \ 1\0V ' - ~ - c

\,

0

F

Distances represent comparative current values as follows: OS = I,,,,, asymmetrical of ac component

CT == It ' l l l " s y m m e t i c a J ~ CB = l lx 'lll> symmetrical EF = Minimum peak current values

OR ::: I,"" asymmetrical; limit to which value approachesNorc: A value slated as closely approaching a designated limitOA I""" asymmetrical; Hmit to which value approachesis considered as cOlnciding with that limit for computation pur

OD = Peak of de component; limit to which value ap poses. An osciJloscope trace shows that the difference is slight inproaches most cases,

Fig, 6-5. Typical curve 01 alternating current wave during offset short-circuit (X /R aboUl 15).

other discontinuities, and damage may also occur to sup

porting structures and hardware.

These phenomena have been extensively studied at out

door test sites m which virtually any type of overhead

conductor operating condition can be duplicated. The

results of many years of sueh research have been made

available to the utility industry by cooperating manufacturers and technical institutes and universities.

Conductor vibration and oscillation may be divided into

three general types;

l . Sway or side swing is the most obvious and simplest

form of conductor movement in an entire span. It

is caused by crosswinds or short-circuit forces,

2. Aeolian vibration is a resonant vibration, It is the leastreadily observed and usually the most damaging type.It is caused by steady crosswinds. The conductors vi

brate in much the same way as any string undertension. Frequencies range from 2 to 200 Hz.

3. Gal/oping or dancing is the movement that sometimes

results when the interrelation of wind direction and

velocity. as well as of moisture and temperature, issuch that the conductor becomes eccentrically glazedor ice-coated, A movement pattern develops in which

the entire span oscillates as a whole or in a few loops,with amplitudes of several feet and at low frequency,

largely in a vertical direction, The envelope of motion usually is an inclined ellipse, Galloping is re

ported to have been seen infrequently even with the

conductors free of ice,

Aeolian vibration and galloping present the most seriousproblems, since either of them may lead to failure of eon

d uctor strands at points of support or at other discontinuities. The most common types of damage are actual

failures of the conductor, the hardware, or componentsof the supports or towers, In addition, there might bedamage and service interruptions caused by phase-te-phaseor phase-to-ground contacts during severe galloping.

Aeolian Vibration ot Conductors

The accepted explanation of the wind-induced phe

nomenon known as aeolian vibration is as follows: When

a comparatively steady wind blows across an overheadconductor under tension, vortices are detached at regular

intervals on die lee side of the conductor-alternatelyfrom the top and bottom portions. The conductor is thus

repeatedly subjected to forces that are alternately im

pressed from above and below. The frequency of these

forces increases with increasing wind velocity and withdecreasing conductor diameter.

I f the frequency of the forces corresponds approxi

mately to the frequency of a mode of resonant vibration

of the span, the conductor will tend to vibrate in many

loops in a vertical plane, As the amplitude of vibration

increases. the vortices tend to be detached in synchronism

6·13




with the vibration to increase the amplitude. The forcesimpressed by the wind on the conductor produce travelingwaves that move away from the points of application of theforces toward the ends of the span. Each wave, i.e., eachcrest and trough, stores part of the energy it receives fromthe wind during the course of its travel, in the form of increased amplitude-the crest becoming higher and the

trough deeper.When a wave reaches the end of an undamped span

and is reflected, neither its amplitude nor the energystored in it is significantly diminished by the reflection.During its subsequent travel, the wave acquires moreenergy and greater amplitude until an equilibrium amplitude is reached where dissipation in the conductor matchesinput energy. At the ends of the span the reflected travelingwaves are superimposed on incoming traveling waves,thereby producing standing waves. The standing-waveloops thus formed have frequencies that are multiples ofthe fundamental frequency of the entire span.

The observed relative absence of vibrations at higher

wind velocities can be attributed in part to wind turbulence. Conductor vibration is usually not observed at windvelocities above 15 mph, although where high tensionsare used and where there are steady winds of up to about30 mph, conductor vibration has been observed. Anotherreason why vibration of significant amplitude does notgenerally occur at high wind velocities is that these causehigh vibration frequencies, and the self-damping orinternal dissipation of energy in a stranded conductorincreases rapidly with frequency.

The tendency of a conductor to vibrate increasesrapidly as conductor tension is increased. Conductor vi-bration is almost never observed at low stringing ten

sions; i.e., less than about 10 to 12 percent of ultimatestrength. Hence, even with dampers, limitations of 25percent final tension and 33 percent of ultimate strengthinitial tension with no ice or wind at the design loadingtemperature were established for controlling aeolian v i b r a ~tion, and are nOw widely accepted.

No exact tension limit can be defined which will assurecomplete self-damping protection, but only rarely hasfatigue damage been observed when tensions have been12 percent of rated strength or less.

In certain areas where local wind turbulence caused bybroken terrain Or trees reduces the power input of wind,somewhat higher tensions have been used on otherwise

unprotected spans without resultant vibration difficulties.In exposed areas with steady winds, however, a few lineswith tensions as low as 11 percent of ultimate have suffered damage.

Fatigue of Conductor Strands·

Close inspection of fatigue failures has shown thatcracks begin at fretted regions where the strands have

$: IEEE Transaction on Power Apparatus and Systems, Vol.PAS-87, No.6, June 1968, pp. 1381,1384, Fricke and Rawlins.

rubbed repeatedly against each other or against an armorod or clamp. Micrographic studies show that the surfacelayer of a strand is severely disturbed by the frettingCracks appear within the disturbed layer and-under thevibration stresses present in the conductor-may penetrate into the undisturbed metal below the fretted region.

The probable explanation of the phenomenon of fret

ting is as follows: Flexing of the conductor at the point ofsupport results in a small amount of movement betweenadjacent strands in the conductor or between strands andadjacent members. At the microscopic level, the contacbetween metal surfaces is not a plane contact but rathera contact between asperities (minute projections). Theintimate contact between asperities, aided by the wipingaction-which removes surface films-results in microscopic welds between the asperities. Further movemenbetween strands, however, breaks these welds or themetal adjacent to the welds. When movements betweenthe strand surfaces are repeated a number of times, manywelds are made and broken, and a disturbed layer is

formed on the strand surface. Debris produced by thefretting can be seen as a fine dust surroun<!jng the frettedarea. Cracks are graduallY opened in the disturbedsurface layer by the forces involved.

Vibration Dampers

Perhaps the first device of any value for reducing 'ibration was the festoon damper, with one or more somewhaloose auxiliary conductors from 4 to 12 ft. long clampedto the tensioned conductor at each side of a suspensionpoint. It was not until about J930 that successful damping control was achieved by the introduction of the Stock

bridge damper, Fig. 6-7. This device consists of two

weights attached rigidly to the ends of a resilient steelcable, which, in turn, is attached to the conductor bymeans of a clamp at the midpoint. Because of the relatively large mass of the damper weights, the steel s u p ~porting damper cable is not stiff enough to force themto follow accurately the motions of the cable clamp, andthis causes flexure of the damper cable, which results inslipping between its strands with consequent dissipatiorof mechanical energy from interstrand friction. If thedamper and conductor span can dissipate energy at agreater rate than that at which the wind imparts it. :hevibration of the span is suppressed to harmless prop·."·rj,,,lS.

The selection of damper sizes and the best placemen

of them on the spans are determined by the tension,weight, and diameter of the conductor and ,he expectedrange of wind velocities. With new efficient damper designs and usual conductor tensions and span lengths, onedamper is installed near one span support point. Forlong spans, additional dampers may be required. Tensionis normally taken as that for "final condition" at about60'F. It has been found that protection from damagingvibration is most evenly balanced over the range of expected frequencies of line vibration when the damper isspaced so it is approximately 70 percent of a free-loop

6-14



Fig. 6-6. Installing a 735,000-voll line of aluminum across

the St . Lawrence River.

length from the fixed end of the span for the highest expected frequency, though this distance may vary with the

design of the damper. Determination of the free-loop

length is as follows:

f = 3.26 V/ d (Eq . 6-2)

and

(Tg/w) y,

FLL=-- - (Eq. 6-3)

2f

where:

f = Frequency of conductor vibration, cycles per

sec

v = Wind velocity, mph

d = Conductor diameter, in

FLL = Free-loop length between amplitude peaksof conductor vibration, ft

g = Acceleration of gravity, 32 .2 ft/sec'

T = Conductor tension, Ib

w = Conductor weight, Ib per ft

Example: Assume a span of 795 kcmil·2617 ACSR at tension of

6250 Ib (20070 of rated strength) ("posed to a steady transverse wind

of up to 10 mph . Substituting values from the conductor tables.

From Eq . 6-2: f = 3.26 X 10/ 1.108 = 29.4 Hz . conductor vibration.


(6250 X 32 .21 1.094) V

From EQ . 6·) : FLL = = 7.J ft . free loop

2 X 29.4

length (from crest 10 crest on the same side of conductor), hence

the spacing would be approximately 0 .70 X 7.3 = S. l l ft from suppon .

Normally. the spacing is increased 0.2 ft to allow fOT approximately

one·half lenglh of Ihe suspension clamp Or insulalOr groove. Allhough

dam pe r spacings usually are given (rom the center or Ihe suspension

clamp or ins ulalOr groove Ihe fixed end is more nearly the point of

tangency nea r the end of (he clamp or groove . A t dead end s. spac ing

is measured from the mouth of the clamp. Precise data in th is reg ard

should be obtttined from th e damper supplier.

Values from Eq. 6-3, modified as noted abo ve , are plot

ted on Fig. 6-8 for a maximum steady wind velocity of

15 mph . Fo r other maximum stead y wind velocitie s ,fa c tor the spacing by multiplying the JS mph di stance by(IS/prefe rred veloc it y in mph). Fig. 6-9 shows a s imilar

solution where armor rod s are used . Armor rods shorten

the end loop by II pe rce nt. When armor rods are used.

they should be of such length that da mpers can be

mounted at proper spacings ju st be yond the rod end s .

Dimensions of Stockbridge-type dampers, weights, and

recommendations as to the number to be used for variousspan lengths are obtainable from the manufacturers.

Other types of vibration dampers have been used in

cluding torsional , impact, spiral, dash-pot, visco-elastic,

and variation s of the Stockbridge with extra weights and

eccentric weights. The most popular system, however, is

the one described.

Spacers and Dampers for Bundled Conductors

Undamped horizontally bundled conductors used on

long-span high-voltage lines with spacers at the customary

250- to 3oo-ft intervals typically vibrate with about half

the amplitude of a single conductor of the same size under

identical conditions. It has been confirmed that the leewardconductor of the pair usually vibrates at greater amplitudethan the windward conductor.

Stockbridge-type dampers are used on the individualconductors of a bundled line . Spacer-dampers, designed to

dissipate vibration energy, are also used frequently. Theyare popular on lines employing three or four subconductors per phase. and provide vibration control as well

as the spacing function.

Figs. 6-10 and 6-11 depict typical types of spacer

dampers. The spherical configuration of the end clampsof spacer-dampers used on EHV lines reduces surface

gradients, thereby avoiding corona.

Fig. 6-7. A Stockbridge damper.

6·15



• •


•'.IotO

,.,""'000

"'"'"'"(1)00

,...

lOOO

lOOO5

lOOO

CO"-iDUcrOIl: OIAMliTEIl: INCHES

(Use rhis graph when armor rods are not employed,)

T = Conductor tension Ib at average temperature.

W Conductor weight Ib/rr.

Fig. 6·8. Spacing between damper and tangent supportcenter to center or to mouth oj dead end. 15 mph maxi·mum vibration· inducing wind velocity assumed.

. . . . •

\

• •

\, \ \1 \\ \ 1\

\ \ \ \ \ \ . \" lI

\ \ . \ \\ \ \ \. \\ \ 1\ \ \. \ 1\ \1\

\ \ \ 1\ \ \ 1\ 1\

\ \ \ \ \ \\ \ 1\ 1\

\ ! I

\ \ \ \ ~ 1\1\,\ \\ ~ \ \ \ ~ \ \1\ ~

(Use this I(raph when armor rods are employed.)

T = Conductor tension Ib at average temperature.

W = Conductor weight lb/ft.

0.2

O A M ~ 1 l : SPACINC -INCHES

",e 1.0

CONO\J(TOP DIAMETeR - INCHES.

,

,000,"0

•'"

..

"'''

"

000

,

,

Fig. 6·9. Spacing between damper and tangem suppOrtcenter 10 center or to mouth oj dead end. 15 mph maximum vibration·inducing wind velocity assumed.

Fig. 6·10. EHV21e bundle/phase spacer·damper. Fig. 6-11. EHV3/e bundle/pnasespacer·damper.

aech chapter 6

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