EFFECT OF CABLE GEOMETRY ON INDUCED ZERO SEQUENCEGROUND CURRENTS WITH HIGH POWER CONVERTERS

Copyright Material IEEEPaper No. PCIC-2006-19

Gary L. SkibinskiRockwell Automation / Allen BradleyStandard Drives R & D6400 W. Enterprise Drive, Mequon, WI 53092(262)-512-7151 gIskibinski@ra.rockwell.com

Bob BrownE. I. DuPont de NemoursDrives Research LabP.O. Box 80840, Wilmington, DERobert. L. Brown-1 @usa.dupont.com

Mark ChristiniAnsoft CorporationFour Station Square, Suite 200Pittsburgh, Pa 15219(570)-265-2544 christini@ansoft.com

Abstract: This paper investigates a field site phenomenon in whicha large zero sequence current was found to circulate in the plantground grid, even though the main supply transformer wasungrounded. The magnetically induced ground current is explainedand seen to be influenced by cable geometry and operatingcharacteristics of non-linear power converters. Finite ElementAnalysis (FEA) models of various cable geometry's are generatedwith their respective Impedance Coupling Matrix models that arefurther embedded into a circuit simulator program to determineinduced ground current magnitude. Simulation results are shown torelate with field ground current measurements. Induced groundcurrent is also analyzed as a function of line current magnitude,linear/non-linear load operation, and phasing arrangement ofparallel cables. The paper concludes using analysis results toformulate application guidelines to minimize the effects of inducedground current phenomenon prior to equipment installation.

1. INTRODUCTION

A primary emphasis of electrochemical applications is on high powerDC rectifier / converter design. The high current magnitudes involvedresult in special considerations given to cable geometry to obtainrequired current balance in parallel rectifier cells [1]. The high currentmagnitudes involved also result in consideration of converterconfigurations that reduce line side current harmonics injected into theutility grid. Most industrial rectifiers usel2-pulse converters that phaseshift one converter by 30 degrees with respect to the other [2,3]. This istypically accomplished by delta/wye and delta/delta secondary windingsthat are floating from ground and that feed two 6 pulse rectifierconverters as in Fig. 1. This paper will show that although the highcurrent secondary windings are not grounded, a zero sequence groundcurrent may flow based on input cable geometry and converter operatingcharacteristics.A need often arises in general industry to also specify a Variable-

Frequency Drive (VFD). VFDs may provide precise torque control(current regulated mode) or provide adjustable voltage - adjustablefrequency (V/Hz mode) to motors, such as fans / pumps, providingoptimum operating conditions and energy savings.

Fig. 2 shows a typical low voltage (< 600 V) VFD topology for acommon dc bus / multi-VFD system. It consists of a utility supplyvoltage feeding Input Cables to a Diode Rectifier Converter and DC BusBus Inductor - Capacitor Filter, which converts the AC voltage to a

r fl IMlain suppilyII |xfmr #1

L____J J ~~~~6Pulse

I----....Mai supply

I xfmr #2_ _ )IiI _ - . ~6 Pulse CommonI/%j _ li _ ~converter #2 dc bus

L cable tray G w

Fig. 1 Typical Electro-Chem 12-pulse converter rectifier section

1-4244-0559-9/06/$20.00 02006 IEEE

A

Main supplyxfmr Common dc bus

t t t (+)

(-)PWMvsI

Fig.2 Typical low voltage common dc bus / inverter section

A" CA CAC 4000A600V3 phase 3 wire

+ ~~~~~~~LR PHASING8" C A C A C A 600V 750MCM CU

80 Mil EPR6 per phase OD-1.28"

4- 18" -*

Fig. 3 Cable duct & phasing with parallel-six 750 MCM wires at 4,000 Arms

Common Bus DC voltage used by multiple inverters. Table 1 shows oneplant used 3 MVA delta/delta and delta/wye sections to help reduce lineharmonics. A DC bus may have several Pulse Width Modulated (PWM)Voltage Source Inverters (VSJ) converting Common DC bus voltage intoadjustable voltage-adjustable frequency. VSI semiconductors connectedto the (+) and (-) DC bus are switched at a high rate (f, = typically 4kHz) to create three phase high-frequency square wave voltage pulseswith sharp transition risetimes that are fed to the AC Output Cables.

PWM VSI's control pulse width on both positive and negative phaseportions of each cycle. Fourier analysis of the pulsed waveform resultsin a fundamental low-frequency fo sinusoidal alternating voltage at theterminals. Output current waveforms are sinusoid due to the highf, rate.

AC Output Cable specification considerations listed below typicallyimplies use of continuous welded aluminum armor type cables [4-7].

- Cost of cable, cost of labor and installation- Reliability under harsh environmental conditions near the load.- Reliability under increased peak electrical stress induced by reflected

wave voltage spikes due to the PWM fast voltage transition rise-times.- Containment of Electro-Magnetic Interference (EMI) ground noise

current between inverter - motor due to electro-static capacitive coupling- Current & voltage balance with parallel cable connections

AC Input Cable specification has a reduced set of considerations.Environmental conditions are closer to standard industrial installations,since input transformer and VFD lineup are closer together, away fromextreme heat and humidity of the load motor. Reflected wave issues onthe utility side do not exist. EMI noise current containment betweeninput transformer and VFD lineup can be alternately achieved withoutshielded cable, with other noise filter techniques [7]. Current andvoltage unbalance considerations are not critical for VFD rectifiers asfor induction motors. Thus, the 3-phase 3-wire cable-duct system ofFig. 3, with individual parallel conductors, is sometimes specified toreduce cabling cost. The aluminum tray in Fig. 2, 3 & 4 is permitted for

Table 1 Common dc Bus system lineup with associated PE ground current

DESCRIPTION

Converter Type:

Internalcable traybond wire

Fig. 4 Physical representation of parallel-six input feeder cable duct system

Transformer:- 13.8 kV / 600V- Pri / Sec

Grounding xfmr- R Gnd ohmsCable tray:

Phase Amps:ABC

Phase Volts:ABC

Line-Line Volts:A-BB-CC-A

A/AFloatingZig-Zag150 QFig. 270 ft.

1200 A1240 A1190 A

,A =40A

349 V350 V350 V

596 V598 V596 V

Diode Diode &PWM Rect.

3 MVA 3 MVA

A/A A/AFloating FloatingZig-Zag Zig-Zag150Q 150QFig. 2 Fig. 245ft. 26ft.

347A 388A337 A 380 A337 A 384 A3616VA =4-8A

366V 366V366V 364V366 V 364 V

598 V 600 V598 V 600 V597 V 599 V

Fig. 5 Zero Sequence current induced in cable tray and plant ground gridPE as result of cable wire phasing causing unbalanced mutual coupling

use as a sole grounding conductor on circuits with ground faultprotection up to 2,000 amps [8].

II. ZERO SEQUENCE CURRENT PHENOMENON

A. Problem Overview

An unexpected amount of zero sequence current was detected in a

plant ground grid of a recently commissioned plant. Measurements ofFig. 2 PE Ground wire, bonding the tray to the common dc bus ac drivesystem Cabinet Ground in Fig. 4, showed a zero sequence groundcurrent of 62 Arms. Measurements of an identical common dc busconfiguration with identical loading at a second site showed muchreduced zero sequence current of 16 Arms in the incoming PE Groundwire. One site difference is in the input power feeder cables. Site #1having a discrete parallel-six cable phasing duct arrangement (Fig. 3 &4) and Site #2 using parallel-six aluminum armor cables in a cable tray.

Fig. 5 represents the input power source feed of Fig. 2 VFD lineup.A lack of conductor phase symmetry results in an unbalance in mutualimpedance coupling between conductors in Fig. 3, so that a net externalmutual magnetic flux coupling to the cable tray ground exists. Aninduced zero sequence current (Izs) circulates in the tray ends and PEcopper grounds to building steel to form a circulating ground loopcurrent, such as a single turn transformer load with the plant groundcompleting the secondary loop.

It is shown the measured 62A zero sequence ground current in the PEground at the tray ends is generated due to unbalanced mutual couplingbetween the parallel circuits. It is also shown the parallel six-conductorcable duct system can never have a cable phasing arrangement toeliminate zero sequence current induced in the ground system.

Zero sequence current magnitude is shown to vary greatly accordingwith conductor phasing arrangement and overall cable geometry. Site #1with bus duct geometry show a (Izs I Iphase load) ratio of 62A / 1500A -

4.l1%. Site #2 with symmetrical armor cable geometry measured a (Izsphase load) ratio of 27A / 1500A 1.82%.

Zero sequence ground current magnitude is also proportional tophase current magnitude. Five other Site #1 common dc busconfigurations in Table 1 have zero sequence ground components, sincethe cable duct is identical in all configurations. Input ac phase current islower, so induced Izs magnitude is also lower. However, (Izs / Iphase load)ratio in Table 1 is similar at 3.4% to 4.8%. The remainder of the paper

explains the reasons for and simulates the Izs ground current.um._eo

(a) 1200A input ph

(b) Resulting induced 62 Arms Izs ground current in50A/div 10 ms /div

rectifier converter

E Ground wire

Fig. 6 Comparison of typical six-pulse phase current waveform with inducedIzs ground current at Field Site #1

FAN PUMP PUMP DRIVE DRIVE

Diode Diode

3 MVA 3 MVA

/\---

I_

TriplePWM Rect.3 MVA

A/YFloatingZig-Zag150 QFig. 232 ft.

520 A509 A502 A

,A =18A

390V395 V395 V

599 V599 V601 V

A/YFloatingZig-Zag150 QFig. 226 ft.

596 A575 A596 A

,A =25A

357 V360V360V

599 V599 V601 V

I PE GNDtoBuildingsteel

---I-

Fig. 6 shows a 1,200 Arms input phase current waveform of acommon dc bus diode six-pulse rectifier. Fig. 6b shows the resultinginduced Izs ground current waveform is "almost" a direct replica of theinput phase current due to magnetic coupling, with exception to thecircled ripple around zero. The Izs waveform is simulated in Section VI.

B. Possible Zero Sequence Sources

Cable conductor phasing and geometry is the largest source of zerosequence current. Other possible sources in Fig. 5 are unbalances in themain and zig-zag grounding transformers and single phase loads.Main supply transformer Viine-line data in Table 1 is within 2 V of 600V

or 0.3% and is balanced and not an issue. However, transformer Viine-neutral is higher than 346V (- 600V/1.732), being 349V (0.9%) to 395V(14%). This may be due to a neutral voltage shift due to line currentunbalance and the fact that all secondary neutrals are capacitivelyfloating with respect to ground.

Zig-zag magnetizing current may not be perfectly balanced resultingin a small net unbalance current. A 150-Q ground resistor limits currentflow through the transformer and therefore cannot account for the 40Arms to 62 Arms Izs ground current.

Single-phase loads in Fig. 5 could be a source of zero sequencecurrent, but only if the source neutral connection was grounded.

Thus, the next section first concentrates on the effect of cableconductor phasing and geometry on magnetically induced Izs groundcurrent. Various parallel conductor-phasing arrangements commonlyused are first investigated, before Fig. 3 cable tray phasing is modeled,simulated and compared to field results for magnetically induced Izsground current.

C. Simplified Analysis of60 Hz Magnetically Induced Voltage

Cable conductor phasing and geometry is the largest source of lowfrequency 60 Hz zero sequence current induced into a ground conductor.Figure 7 below shows an un-symmetrical and symmetric arrangement ofphase conductors (a, b, c) with respect to a ground conductor (G,).Eq. (1) defines the voltage induced in ground conductor G0 that is

parallel to a,b,c in Fig. 7a geometry [9]. V, is the voltage induced inconductor G, per mile, Ia is 60 Hz rms positive sequence phase currentflowing in a,b,c, Dax is distance between phase a and G,. Likewise, Db,and DC are defined similarly. Dax, Db, and D, can be defined in anyconsistent units of measure. The first term is positive for phase rotationa, b, c and negative for a, c, b.

v=0.2794 Ia -L2 LogI0 Db KJLoiCDbx+±Dcx

Dax )

Eq (1)

Examination of Eq.(1) shows induced ground conductor voltage V, is

directly proportional to Ia. Also, unequal Dax, Db,, D, distances of Fig.7a maximizes ground voltage. Equal Dax, DbX, D, distances of Fig. 7b

minimizes ground voltage, since the 1st Log term is zero and 2nd Logterm is j 0.15, resulting in a smaller V=j 0.042 Ia. Thus, a seeminglyperfect symmetric arrangement may have induced voltages on parallelconductor G,. While induced voltage may exist on conductor Gx, its

current flow depends on the external loop impedance connecting both

ends. Consider G, as the aluminum cable tray in Fig. 4 & 5. If the return

loop path of Fig. 5 is the earth's resistivity, then cable tray current maybe low depending on soil conditions. Foot-ohms for swampy ground is

32, average soil is 330, while dry earth is 3280 [ref. 9, p.580].

\ Db x ( <

ax Iscx

(a) Un-symmetric ground arrangement (b) Symmetric ground arrangementFig. 7 Typical spacing of phase & ground conductors G,

However, if the cable tray ends are bonded to a building column on eachend, which in addition has welded copper conductors connecting multi-point building steel, then the loop resistance is very small and evensmall G, induced voltage may induce a large G, current, that again isdirectly proportional to Ia Pro's and con's of single-point ground vs.multi-point ground vs. safety requirements has been a topic of manyPCIC conferences.In the following sections, this paper will further investigate zero

sequence current with regard to typical cable geometry factors and phasecurrent Ia magnitude. A detailed zero sequence current analysis usingFEA models of the more complex multiple parallel conductor cable trayproblem is done. The paper finally uncovers an induced ground currentproblem with high current non-linear converters which may even use thebest commercially available perfectly symmetric cable geometries.

III. INDUCED ZERO SEQUENCE CURRENT ANALYSIS

Reference literature exists on the phenomenon of zero sequencecurrent magnetically induced in a ground grid as a result of parallelcable phasing arrangement. Reference [10,1966] states "zero sequencecirculating currents will flow in parallel un-transposed multi-circuitlines". The parallel-6 multi-circuit bus duct under discussion is of shortlength, 80 ft. to 100 ft., making a transposition solution complicated.Thus, zero sequence ground currents induced in cable bus duct systemsis not a new phenomenon; it is a question of what factors affect itsmagnitude. Reference [11,1991] further substantiated an unexplainedzero sequence current phenomenon in high voltage power systemtransmission lines using Electro-Magnetic Transients Program (EMTP)simulation runs, mathematical analysis and gathering of field data.

This section first summarizes findings on zero sequence currentinduced with parallel-two and parallel-four circuits'[1 1].Thisinformation is then extended to explain the zero sequence phenomenonin the parallel-6 cable bus duct circuit under discussion.

A. Parallel Twin Circuits in Nine-Duct System

Reference [11] analysis shows large I, can be generated in cable ductsystems due to unbalanced mutual coupling between parallel circuits.Zero sequence current is the vector sum of the zero sequence current inits own circuit and the zero sequence currents induced from the otherparallel circuits. Table 2 shows zero sequence currents vary greatlyaccording to cable phasing arrangement, for a circuit load of 500 Arms.

Low ReataneLR)PgLg&ointS alCase 1 and Case 2 are inversely arranged phasing arrangements calledLow Reactance Phasing or LR Phasing. This bus duct method iscommonly used in industry. Phases A &C are inversely arranged aroundthe center Phase B. Phase A, B and C are point symmetric around acentral midpoint between Phase B parallel conductors, "bl" and "b2", inthe empty conduit hole. Conductor point symmetry is the first importantcriterion to minimize circulating zero sequence currents, while axialsymmetry is the second. Phase B is axial symmetric for any vertical linedrawn through Case 1 bus duct, but Phase A and C are not axialsymmetric. Insignificant I * is generated in Case 1, since it is pointsymmetric but not axial symmetric.

Case 2 is similar with the empty conduit between circuits removedand closer phase distance separation. Table 2 shows a slightly higher I,= 0.03 A due to stronger interaction of unbalanced mutual impedance's.

Supr un le etPaig &AiaS etialCase 3 and Case 4 are defined as Super-Bundle Phasing or SB Phasing,since Phase A, B and C conductors of each parallel circuit are bundledvertically. Perfect vertical axial symmetry exists in Case 3, so "no" I * isgenerated. Case 4 is similar but with conduit between circuits removed.

Case 5 and Case 6 are defined as Triangular Conductor Phasing, sinceCircuit #1 and Circuit #2 phase conductors are arranged in a triangle.Phasing is point symmetric with respect to the midpoint center of allcables. Insignificant I * is generated in Table 2 due to point symmetry.Case 6 is similar with conduit between circuits removed.

Case 8, Case 9 and Case 10 are asymmetrical conductor phasingarrangements. Symmetrical conditions require phase conductors to bemirror images, (e.g. "'a1"' & "a2") on opposite sides of any vertical,horizontal or diagonal line drawn. Asymmetric cases generate extremelylarge I, in Table 2. Case 8, Case 9 and Case 10 respectively have I,values of 27%, 23% and 37%, relative to the 500 A circuit load.

Table 2 Parallel Two Circuit 1,, for 500 Amp circuit load

PHASING Iz A Ieg PHASING I A IdegPOINT SYMMETRIC: LR PHASING

0 0.01£- 28.0

o Do o.oiL 52.0

2 POINT SYMMETRIC: LR PHASINI

QOG 0.03£ 168.9Ibli b21 )81-91 0.03£ -14.4

AXIAL SYMMETRIC: SB PHASING

00AXIAL SYMMETRIC: SB PHASING

bl!b0 o310

PP TS MMETRIC: TRI PHASING

~o.olL 7.4OOQ(30.01L - 175.7

6: POINT SYMMETRI: TRI PHASING

/-,\Io30.01£-7.4o.olL@0-1 - 175.7

,ASE 7 AXIAL SYMMETRIC:

o.o1£ 155.4

OOQ@o0 01£- 18.5

,ASE 8 ASYMMETRIC

i3@135 L 144.0

135 L- 36.0I, - 27% Of load

,ASE 9 ASYMMETRIC

QO1g2 115 L- 21.4a0° (@ 115 LC 158.6

(i)0 (i t 23% Of load,ASE 10: ASYMMETRIC

184 £- 141.7184 £- 38.3tz~37% Of load

ASE 11: SYMMETRIC: PLANAR PHASING

0

0

Case ]] is a planar phase arrangement, as was Case 10 (Is =37%), but isaxial symmetric to a vertical line between cl and c2. Thus, I * is 0 %.

Math analysis of the Parallel Two Circuit cases proved the requiredcondition to insure no I, is to have all mutual impedance's betweenconductors balanced [11]. Proper conductor phasing arrangements canaccomplish this. A zero I * condition corresponds to point or axialsymmetry of the six-conductors in the two parallel circuits and indicatesan asymmetrical arrangement of cable is the cause of zero-sequencecurrent generation. In other words, the mutual impedance unbalancegenerates the zero sequence circulating current in the twin circuit.

Case 3 has four parallel circuit combinations of Triangular ConductorPhasing so all conductors are point symmetric with respect to a centralmid-point between "c2" and "c3". I, - 79A for a 500A circuit load,resulting in a (Is /1load) ratio of 16 %.

TRK LR Codco PhasetrialCase 4 combines (L1 & L2) parallel circuits using Triangular ConductorPhasing with (L3 & L4) parallel circuits using Low Reactance ConductorPhasing. Conductor phasing is asymmetric with respect to all fourcircuits. LR phased lines L3 & L4 have I, - 13A resulting in a (Is/Iload)ratio of 2.6 %. TR phased lines L1 & L2 have 1, - 61A resulting in a(Is/IIoad) ratio of 12 %.

Lo eCtac (_R Phfflen Pit S calCase 2 arrangement is point symmetric with respect to the midpointcenter (between "b2" & "b3") of the 12 conductors. Izs - 20 A for a500A circuit load resulting in a (IzS1Iload) ratio of 4 %.

Table 3 Parallel Four-Circuit Izs for 500 Amp circuit load eachL: a1, b1, c1 L IL2: a2, b2, c2 L2/L3: a3, b3, c3L3L4: a4I b4, c4 4

PHASING lzs A degCASE 1 AXIAL SYMMETRIC: SB PHASING

CAE2POINT SYMMETRIC: LR PHASINGRoute A Route B

Ll: 21 L-132.1

,..~~ O L2: 21 £ 47.9~~K~~ L3: 20 L 47.9

. ..... L4: 20 L 132.1LR LR

CASE 2:POINT SYMMETRIC: TRI PHASING

Route RouTR L: 78 L£312.1

Rote 8 TR L4 78 /3.

CS ASgYMMETRI-C: TR & LR PHASING

LR L2: 61 37.9

%ffi(ig£2)L2: 61 £-142.1(C9 L3: 13 L 47.9

Ruei2 TR113214: 13 £ 37.6

CASE 3r POINT SYMMETRIC: SB PHASINGRouute B L1 20 L.1321

Ob (@'@ / L4: 20 L -132.1SB SB

B. Parallel Quad Circuits in Twelve Duct System

Four parallel circuits in a nine-duct configuration are reviewed. Thisconfiguration is close in kind to Field Site #1. Table 3 shows generalcable arrangements widely used. Math expressions to define thebalanced condition are too complex, so simulation was the main tool.

Supe7rBundle 'SB) Phqsix Axial & Point S metricalCase and Case 5 are Super-Bundle Phasing arrangements. Case isaxial symmetric to a vertical line drawn between "a1" & "a2". Case 5 ispoint symmetric to a central midpoint between "b2" & "b3". Both cases

have Izs -20 A for a 500A circuit load, resulting in (I,Ihoad) ratio of 4 %.

Zero sequence currents vary with conductor phasing arrangement.Results show Izs generated is due to unbalanced mutual couplingbetween the four parallel circuits. Two parallel cables, Route A Cable#1 & #2, can be arrangedfor zero Izs. However, with four circuits, zerosequence current induced from Route B Cable #3 & #4 is added toCable #1 & #2. Thus, unlike parallel two-circuit arrangements, Izscirculating current cannot be reduced to zero in parallel four-circuits,even with point or axial symmetry conductor configurations. Resultsshow a minimum (Izs /1load) ratio of 4 % can be expected. Conductortransposition is required to reduce Izs in a four-circuit cable duct system.

a - - - .- -11 s_T. - 11 A \sl- 1 11Tl - I ._,11rASEF 1

C. Extension to Parallel Six-Circuits in Eighteen Duct System

Fig. 3 six-circuit cable duct is conceptually similar to Table 2 four- S4circuit Case 2, but with two additional parallel circuits. Point symmetryexists like the parallel four-circuit Case 2, evidenced by the traymidpoint dot between Fig. 3 Phase B conductors. LR Phasing exists likefour-circuit Case 2, evidenced by the vertical rectangular boxes in Fig.2 with alternate CBA and ABC. Thus, four-circuit Case 2 Izs magnitudes 2- 4 % of load can be expected in the six-circuit case due to similargeometry. Measured field site Izs data in Table 1 was 3.4% to 4.8% oftotal phase current load for the six-circuit cable tray duct.

D. Parallel Five-Circuits NEC Code Recommendation 4

The NEC Code prior to 1996 required parallel single conductors to beinstalled in a single layer. However, in lieu of previous analysis a single 5layer approach is not desirable due to unequal phase reactance, voltageunbalance and induced Izs current. Effects of cable geometry on theseundesirable characteristics inspired a 1996 NEC code exception to allow 6the more technically desirable circuit group triangular bundling optionof Fig. 8 [11]. Air space between circuits provides thermal cooling aswell as increased spacing to decrease magnetic coupling betweenparallel circuits. Applying Ampere's law to a circuit group with acircular magnetic path length dl shows minimized external magneticfield Hto other parallel circuits, since N = 1 and ia+ ib + ic 0.

fH.dl=ZN.I (2)

NEC Code 1996: 318-8(e): Single conductors where any of the single conductorsinstalled in a ladder or ventilated trough cable trays are Nos. 1/0 through 4/0, allsingle conductors shall be installed in a single layer.Exception: where conductors are installed in accordance with Section 318-11(b)(4), the conductors that are bound together (triangular configuration) to compriseeach circuit group shall be permitted to be installed in other than a single layer.

Table 4 FEA Parametric Results for Balanced Current Conditions

,etup Tray PhasingMaterial

Aluminum aaaaaabbbbbbcccccc

Aluminum acacacbbbbbbcacaca

Aluminum acbacbbacbaccbacba

Steel aaaaaabbbbbbcccccc

Steel acacacbbbbbbcacaca

Steel acbacbbacbaccbacba

F1lux L ine s

_ 1. 9380e-004_ 1. 5795e-004_ 1. 22I0e-004l|8. 6247e-005

...........i 5. 0398e-0051. 4549e-005

-2.130Ole-005

-9.2999e-005 .._-1. 2885e-004-|-1. 6470e-004 ..

I...........

I-real[A]

1.66

I_imag[A]

3.35

I_cmpxmag[A]

2332.8

PowerLoss [W]

152.0

::Airspace d

0 0 0,,

v . '- --

Fig. 8 Circuit group triangular bundling option in cable tray

IV. FINITE ELEMENT ANALYSIS OF CABLING MODELS

A. Bus Duct FEA Model Setup

Initial investigation into Fig. 3 Field Site #1 bus duct cablingphenomenon first focused on use of FEA tools [13] to study the electro-magnetic effects under different cable phasing arrangements and fordifferent cable tray material, being aluminum or steel. Initial studyassumed balanced line current conditions, results in later sections are

given for unbalanced cable currents.

B. Condition 1: Balanced Current - Different Cable Phasing

Model setup consisted of six-750 MCM cables per phase, spaced inthree rows, inside a 10" x 18" cable duct with 1/8" thick casing andanalyzed as a two-dimensional eddy current problem [13]. Source setupconsisted of a 60 Hz 500 Amp magnitude intentionally forced on eachstranded conductor in Phase A with 500 A at 1200 for Phase B and 500A magnitude at 2400 phase angle for Phase C.Parametric setup of different phasing arrangements, tray permeability,j=stee 500 & !taluminum= 1] and tray conductivity [Gsteel= 5*106 &

aluminum 1* 107 Siemens/m] is shown in Table 4 for the balanced study.

Fig. 9 Flux line plot into cable tray for Setup 2, similar to Field Site #1

Fig. 10 Current density vector J plot [A/m2] for cable duct Setup I phasing

Fig. 9 shows typical flux lines [Wb/meter] for Setup 2. The depth ofthe cable tray is into the paper. The flux lines cutting the cable trayinduce 60 Hz eddy currents into the aluminum tray. A Setup 1 currentdensity vector plot at t= 0 in Fig. 10 shows a large current densityconcentration at each 500 amp wire, as expected, along with a highcurrent density in the top and bottom of the tray due to circulating eddycurrents induced in the tray. Ihmpxmag in Table 4 and Fig. 10 is an inducedcurrent that circulates in the tray but does not leave the tray ends.Icmpxmag cannot be measured in the tray material, however it does inducepower loss. From Fig. 10, - 1/2 of Icmpxmag 2,332 A magnitude fromTable 4 Setup flows in the top tray with the arrow direction shown andlikewise l/2 of Icmpxmag flows in the bottom tray, with the sides of thetray showing a much smaller current density concentration.

1.55 3.34 509.9 83.4

1.50 -1.06 530.2 83.8

0.31 -0.04 1038.9 121.2

0.07 0.57 252.4 82.7

-0.256 0.306 276 83.4

i I

A post processing macro was written to calculate the currents andpower loss tabulated in Table 4 for the various setup conditions. Currentis obtained by integrating the Current Density Vector J over the crosssection area, with the tray cable 1/8" thickness playing an importantdimensional role. Table 4 circulating eddy current (hmpxmag) does flowin the tray, resulting in eddy-current power loss [W], regardless ofphasing arrangement and tray material. Table 4 shows for similarphasing arrangements, steel hmpxmag values are -1/2 that for analuminum tray, due to the conductivity of steel is also 1l/2 that ofaluminum. However, since magnetic losses of steel are higher, theinduced power loss values are comparable. The best phase arrangementto reduce hmpxmag and eddy current induced power loss is the LR phasingarrangement acacac, bbbbbb, cacaca for Setup 5 steel and Setup2_aluminum as was used in Field Site #1.

Net current flow through the tray and out the tray ends that anammeter would read is determined by Eq (3). Using Table 4 Setup 2values results in a potentially measurable tray current of 3.6 A.

2 2

icurrent thru and out of tray = I real + I _ imag (3)

Thus, for the parallel six cable duct system with a 3,000 A balancedthree phase load current with 500A/conductor /phase there is a net 3.6 A(0.12%) current out of the tray ends and into the ground grid from thisanalysis. The next step is to purposely unbalance the phase current andobserve the effect.

C. Condition 2: Unbalanced Current- Setup 2 Cable Phasing

FEA parametric results for a parallel six-cable duct system with phasearrangement acacac, bbbbbb, cacaca in aluminum tray [Setup 2] wasrerun with various unbalanced currents from 20 A to 100 A in Phase Aconductors. Case 1 is the balanced condition as in Table 4. Case 2example shows an unbalanced current of 20 A in each of six Phase Aconductors resulting in a Al of 120 A. A post processor routine similarto previous section calculated I real, I imag, I cmpxmag and eddycurrent power loss in the tray. I tray is calculated using Eq 3.

One result of Table 5 shows induced eddy current loss in thetray does not increase appreciably when a phase unbalance of upto 600 Amps occurs. The more important result indicates anyphase current unbalance appears to return through the cable trayand out the ends of the tray into the ground grid, sincemeasurable I tray value calculated is equal to the total phaseunbalance current. The FEA simulation assumes a perfect zeroresistance return for current outside of the tray.

Table 5 FEA Parametric Results for Setup 2 with Unbalanced Current

magnitude or induced Izs cable tray current magnitude seen in the realworld, but provides a valuable tool for further Izs computation work.

Table 6 Inductive coupling matrix model for Setup 2 parallel-six bus duct

duct cll c12 c13 c14 c15 c16 c21 c22 c23 c24 c25 c26 c31 c32 c33 c34 c35 c36duct 1.00 .78 .79 .80 80 .80 .78 .79 .80 .80 .80 .80 .79 .78 .79 .80 80 .80 .78cll .78 1.00 .91 .86 .82 .79 .75 .91 .89 .86 .82 .79 .76 .86 .86 .84 .81 .78 .75c12 .79 .91 1.00 .91 .86 .82 .79 .89 .91 .89 .86 .83 .79 .86 .87 .86 .84 .81 .78c13 .80 .86 .91 1.00 .91 .86 .82 .86 .89 .91 .89 .86 .83 .84 .86 .87 .86 .84 .81c14 .80 .82 .86 .91 1.00 .91 .86 .83 .86 .89 .91 .89 .86 .81 .84 .86 .87 .86 .84c15 .80 .79 .82 .86 .91 1.00 .91 .79 .83 .86 .89 .91 .89 .78 .81 .84 .86 .87 .86c16 .78 .75 .79 .82 .86 .91 1.00 .76 .79 .82 .86 .89 .91 .75 .78 .81 .84 .86 .86c21 .79 .91 .89 .86 .83 .79 .76 1.00 .91 .87 .83 .80 .77 .91 .89 .86 .83 .79 .76c22 .80 .89 .91 .89 .86 .83 .79 .91 1.00 .91 .87 .83 .80 .89 .91 .89 .86 .83 .79c23 .80 .86 .89 .91 .89 .86 .82 .87 .91 1.00 .91 .87 .83 .86 .89 .91 .89 .86 .82c24 .80 .82 .86 .89 .91 .89 .86 .83 .87 .91 1.00 .91 .87 .82 .86 .89 .91 .89 .86c25 .80 .79 .83 .86 .89 .91 .89 .80 .83 .87 .91 1.00 .91 .79 .83 .86 .89 .91 .89c26 .79 .76 .79 .83 .86 .89 .91 .77 .80 .83 .87 .91 1.00 .76 .79 .83 .86 .89 .91c31 .78 .86 .86 .84 .81 .78 .75 .91 .89 .86 .82 .79 .76 1.00 .91 .86 .82 .79 .75c32 .79 .86 .87 .86 .84 .81 .78 .89 .91 .89 .86 .83 .79 .91 1.00 .91 .86 .82 .79c33 .80 .84 .86 .87 .86 .84 .81 .86 .89 .91 .89 .86 .83 .86 .91 1.00 .91 .86 .82c34 .80 .81 .84 .86 .87 .86 .84 .83 .86 .89 .91 .89 .86 .82 .86 .91 1.00 .91 .86c35 .80 .78 .81 .84 .86 .87 .86 .79 .83 .86 .89 .91 .89 .79 .82 .86 .91 1.00 .91c36 .78 .75 .78 .81 .84 .86 .86 .76 .79 .82 .86 .89 .91 .75 .79 .82 .86 .91 1.00

Section III A & B has shown that Izs magnitude generated is due tothe effect of unsymmetrical cable geometry creating an unbalancedmutual coupling between parallel circuits. The final task of the 2Dmagneto-static FEA solver is to generate this inductive coupling matrixmodel. The inductive coupling matrix in Table 6 is typical for the 18-wire bus duct model, where the mutual inductive coupling from eachwire to the other 17 wires is modeled, as well as each wire to the cabletray bus duct. Coupling nomenclature is Cij where i = 1...3 for phase A,B, C wires andj = 1... 6 for each of the six parallel wires per phase.

V. CIRCUIT SIMULATION MODELS FOR lzs COMPUTATION

Fig. 11 shows the circuit simulator [14] with input 6-pulse converterwires simulated with the inductance-coupling matrix model specific tothe cable geometry chosen that was generated by the 2D Finite ElementAnalysis (FEA) solver. The inductance coupling matrix sub-circuit forthe bus duct model is shown with the 6 parallel wire connections madeto the sub-circuit. The unbalanced mutual impedance coupling betweenwires and wires to tray ground are accounted for in the circuit simulatormodel to determine the resulting ground current in the tray. Induced Izsground current into the loop formed with the cable tray with the trayends grounded in Fig. 4 & Fig. 5 can thus be simulated. Circuitparameter R Ground can be adjusted to physically represent the groundgrid impedance possible in Fig. 5. Note Fig. 11 simulated Izs groundcurrent waveform is similar to Fig. 6(b) field site waveform.

VI. SIMULATION RESULTS OF INDUCED Izs

Case PhA[A]

1

2

34

500520540600

PhB[A]

500500500500

PhC

[A]

500500500500

Al I_trayEq (2)

[A] [A]

0

120240600

3.6117.5239.9601.0

I-real I_imag I_cmpxmag I

[A] [A] [A]

1.52-116.3-239.7-600.5

3.60-16.4-10.1-24.9

521.6564.8983.61207.4

D. Inductive Coupling Matrix Modelfor Bus Duct and Arm^Cable Geometry's

The 2D Magneto-static FEA solver of the parallel six bus duct 1

has shown that if we force a balanced conductor current situationthere is negligible induced cable tray currents flowing out of theAlso, if a phase current unbalance condition is enforced, then an inIzs cable tray current flowing out of the tray ends will occur. In gethe 2D FEA program alone cannot solve for unbalance phase c

Power The FEA impedance-coupling matrix embedded into the circuitLoss simulator allows induced Izs investigations under a variety of physical[WI different situations. The effect of bus duct vs. armor cable geometry,83.6 armor type being aluminum vs. galvanized steel, and cable tray height86.1 above the floor ground plane can be studied. Magnetic flux extending8718 outside the cable tray can be plotted to determine interference with96.5 nearby adjoining circuits. Induced Izs with a Non-Linear Converter

Rectifier Diode Load (Fig. 11), a resistance - inductance Linear Load orior PWM Rectifier Converter Load can also be determined.

A. Function ofCable Geometrymodeli, then Field Site #1 of Section II A using Fig. 3 LR phasing cable duct systemo tray. was simulated in Fig. 12.iducedmneral,,urrent

11

0 - a.- A-- -- --

R_round J

roundCurrent Iurr& aae

..5

91 9 41t |i.

-0 k. -~ -'. M m. M~ 2'. 0~ - ~ I1.1 .5

Mn~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~!k.-0!.3\3~m.0 5.11

Fi.1 ici iuao wt aacd3paesure meac opigmarxmdlo nu iest us ovre arsliggon urn

Field Site #2 using parallel armor cables consisting of 3 symmetricalphase wires, 3 symmetrical copper grounds and grounded aluminumarmor was simulated in Fig. 13. Phase conductors A, B, C are triangularsymmetrical with respect to each other so mutual coupling impedance'sare nearly identical within a cable. Thus, there is less chance of phaseunbalance and induced zero sequence components into ground. Phase-to-ground mutual coupling impedance, while not perfectly symmetrical,is at least identical. The 60 Hz stray magnetic fields generated internallyto the cable are not attenuated by the external aluminum armor, so that a

second parallel circuit group cable placed next to it may have some

mutual coupling from the first cable and vice-versa. However, due tobetter symmetry and need for inherent cable tray spacing for thermalreasons, induced coupling between parallel conductor circuits should besubstantially lower. NEC code allows aluminum armor to be used as a

ground conductor, so internal copper wire grounds are smaller [8]. Agalvanized steel interlocking armor would attenuate 60 Hz mutualcoupling fields between parallel circuit group cables. However, steelarmor is not recognized as a grounding conductor and full size copper

grounds must be used inside [8].

B. Function ofTray to Ground Plane Height

Both geometry systems were simulated with a return ground plane at 1

ft. and 8 ft. distances below the cable tray. The 8-foot spacingrepresents the floor ground plane below the tray as in Fig. 4. However,Fig. 4 also shows many parallel conduits 1 ft below the cable tray whichhave measurable induced Izs currents in the conduit wall. Conduitsbonded to the drive cabinet ac power feed section showed replicawaveforms of the 41 Arms Izs waveform in the PE ground wire at thecabinet entry. The three closet conduits to the entry each measured 5 to7 Arms common mode current on site. Conduits farthest from the MCCcable tray ground wire ground had lower Izs current measured at < 1

Arms each. For this reason, simulations were carried out with a conduitground plane at 1-foot spacing below the cable tray in Fig 12 & 13.

A rC* A rC:: A:: CLR

Phasing B:B:B B B B

C A C: A: C: A

$8 ft or 1 ft spacing

Fig. 12 LRphased cable bus duct system with parallel six 750 MCM wires

i

8 ft pr 1 ft sacing

Fig. 13 Symmetrical armor cable system with parallel six cables

C. Function ofArmor Type - Aluminum vs. Galvanized Steel

The 60 Hz stray magnetic fields from each bundled armor cable may

induce a current into an adjoining circuit group cable if the armor is non-magnetic aluminum material. Thus, simulations were done for currentinduced in the aluminum tray ground if the armor was galvanized steel.

R_noutral

IIi

R Load

D. External Magnetic Flux Plots

Fig. 14 shows magnetic flux lines spray outside the internal duct spaceand into the aluminum tray duct due to asymmetric geometry of the 18cable duct wires. The aluminum tray duct acts as a one-turn transformerconductor with the PE ground return wires and ground grid 8-footbelow, completing the secondary loop as in Fig. 5. The inductivecoupling path exists even though the system-input transformer isfloating with respect to ground. The higher the load current the higherthe stray magnetic field and higher induced ground current. Flux densitynear the tray is 2.3 e-5 Wb/m2 and is seen to extend at least 4 feet belowthe tray. This explains why parallel steel conduits 1 foot below the traywere also found to have induced Izs current flowing in them.

Fig. 15 shows a magnetic field plot for a parallel 6-cable geometry,consisting of symmetric A_B C triangular phased aluminum armorcable with 3 symmetric grounds. The triangular bundle has less externalmagnetic flux lines into the aluminum tray ground. Each aluminumarmor has low flux concentration, thereby allowing coupling flux to theadjacent wires in nearby cables as in Fig. 17. Flux density near the trayis 2.9 e-4 Wb/M2, a 12x reduction factor compared to Fig. 14.

Fig. 16 shows a magnetic field plot for a parallel 6-cable geometry,consisting of symmetric A B C triangular phased galvanized steelarmor cable with 3 symmetric grounds. There are minimal externalmagnetic flux lines into the aluminum tray ground. Each steel armor hasa large flux concentration, thereby reducing coupling flux to the adjacentwires in nearby cables as in Fig. 17. Flux density near the tray is 1.5 e-7Wb/M2, a JOOx reduction factor compared to Fig. 14.

Plotl

_ 1.9385e-0041. 1448e-0046.760o3e-0053.9923e-0052.3576e-0051.3923e-0058.222 le-004.8555e-0062.8674e-0061.6933e-00 61.0000e-006

Phasing of 18 Cables:ACACACBBBBBBCACACA

PE Ground Wire

Fig. 14 Magnetic flux plot for parallel-six 750 MCM wire laid in aluminumcable tray with LR phasing connected to a 2,200 Adc rectifier

Flux Lines ..

_ 8. 8707e-0047. 3913e-0045. 9119e-0044. 4325e-0042. 953le-0041. 4737e-004

-5. 646le-007-1. 4850e-004-2. 9644e-004- 4 443 8e-004

Flux Lines ...

4.4598e-0073.7235e-0072.9872e-0072.2510e-0071.5147e-0077.7844e-0084.2178e-009-6.9409e-008-1.43O4e-0072.1666e-007-2.9029e-007

Fig. 16 Magnetic flux plot for parallel-six galvanized steel armor cable using3-symmetrical grounds & A B C triangle phasing with 2,200 Adc rectifier

Fig..17.Magnetic.flux detail: Aluminum(left) vs. galvanized steel armor.cable

Fig.17Magneticfluxdetail:Aluminum(left)vs.galvanizedsteelarmorcable~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~.

E. Operation with Non-Linear Converter Diode Rectifier Load

This section shows that diode rectifier line currents conduct only twophases at a time and that even if one obtained a "perfectly symmetrical"3-phase cable geometry, cable operation itself would be unsymmetricaland always results in a unique Izs ground current characteristic.

Operation with a non-linear diode front-end converter hascharacteristic non-continuous 6-pulse cable line currents. Fig. 18 showsIA is conducting for two positive 600 intervals, zero conduction for one600 span, conducting for two negative 600 intervals, zero conduction forone 600 span and then repeats. Observation of Fig. 18 shows only two ofthree phase currents are conducting and carrying 2,220 Adc load currentover any given 600 interval. Further investigation of Fig.18 revealsinduced Izs peaks in a 600 span are highest (12 Apk to 18 Apk) whenphase B and either phase A or phase C is conducting. Induced Izs peaksin a 600 span are lowest (7 Apk) when phase A and phase C isconducting.

4k DIODE CONDUCTION SEQUENCE B 'BI

BIB.EC EC!IIA IB.1 ISC RR SoUreaJi

3k r; a r n ReSorc. [A]

01k 1a IV-XA TJ-

~1k

-3k

_4k

IB

50m 55 6m 5m 7OT 75m 80T 851900m 95mF 01 t [sI

Fig. 18 IA, IB, IC [1 kA/div] and Izs currents [IOA/div] for 2,250 Adc diodeconverter using parallel-six 750 MCM wire in cable tray bus duct

Fig. 15 Magnetic flux plot for parallel-six aluminum armor cable using 3symmetrical grounds & A_B C triangular phasing with 2,200 Adc rectifier

I

-11.,

j

1 2 3 45 6

PointSymmetricrI D W :(9:D

& A :C A C>:A C

LRPhasing :0 :®0 O :I:

8ft orl ft spacing

Fig. 19 Equivalent cable duct operation when phase A and C conduct withdiode rectifier load, operation is point symmetric and LRphased

1 2 3 45 6

IA A ANot : A

Point I0 O:~1O 0Symmetric B :B. B. B B B

®: :®: 4:®A A A

BftorlIftspacing

Fig. 20 Equivalent cable duct operation when phase B and A conduct withdiode rectifier load, operation is NOT point symmetric

Fig. 19 representation provides a rationale for smaller induced Izsduring IA and IC conduction periods. Assume IA is carrying a 2,200 Adcload with six parallel cables from the transformer (+), while IC returnsthe same 2,200 ADC load back to the transformer secondary (-). Underthis condition, cable operation is perfectly point symmetric around themidpoint of the bus duct. Also LR Inverse Phasing for the six sections ismaintained. Phase A and Phase C cables appear to be inter-stitched like atwisted pair to minimize external magnetic flux being induced in thealuminum cable tray.

Fig. 20 shows why induced Izs peaks in a 600 span are highest (12Apk to 18 Apk) when phase B and either phase C or phase A (A isshown in Fig. 20) is conducting. Under this condition, cabling in theduct carrying the 2,200 amp load current is NOT perfectly pointsymmetric around the midpoint of the bus duct. External magnetic fluxcan thus be being induced in the aluminum cable tray.

Thus, Fig. 18, 19 and 20 provide an explanation and correlation withthe Izs ground current waveform measured in the PB ground wire in Fig.6(b). Simulated Izs current magnitude as 00 of peak line current is 18Apk /2200 Apk or 0.820o which is somewhat lower than the measured 300 value, but acceptable considering the complexity and multiplicity ofground tie points in a real plant.

F. Operation with Linear Load

Fig. 21 shows induced Izs still exists even for a wye-connected 3- phaseresistance-inductance [0.05 Q, 525 ~th] balanced load with feeder wiresinstalled in the cable duct system. Simulated Izs current magnitude as Ooof peak line current is 10 Apk /1,500 Apk or 0.70o

2k

15k~~~~~~~~~~~~Alk

-05k

'lk

1 5Sk

-2k

-245k0 5m. ICO 1I5m 20m 25m 30m 35m 40m 45m50 mt []

Fig. 21 IA,'B,C 1500A/divl & lzs current 15A/divl for 1,500 Apk balancedlinear R-L load using parallel-six 750 MCM wire in cable tray bus duct

C. Operation on PWMRectifier Converter Load

Active front-end converters, using IGBT semiconductors switched at a 2kHz rate in lieu of Fig. 11I diodes, was also used at Site #1 and Site #2 asa PWM rectifier in Table 1 to a common dc bus line-up. Utility phasecurrent wv~xeforms are more sinusoidal w~ith all three phases conductingcurrent like Fig. 21 linear load case. Circuit simulation was not done.Fig. 22 does show induced Izs in cable tray PB ground wire measured atSite #1 is a near sinusoidal waveform with 2 kHz switching ripple.

Fig. 22 Measured Izs current 150A/divl for 1,250 Arms PWM rectifier loadusing parallel-six 750 MCM wire in cable tray bus duct. Izs- 30o of load

VII. FEA I SIMULATION RESULTS vs. lzs FIELD SITE DATA

A. Measured I, Field Site Data

Fig. 23 I,, measured in Fig. 4 cable tray PE bond wire for 950 Arms inputrectifier load but different input cable geometry, 18 conductor duct *i 32Arms vs. parallel 11 6 Al armor in tray w 16 Arms, [10 ins, 50A per divi

20

10 M A

0

10 r

20

Fig. 24 Simulated waveform ofFig. 23a Field Site #1 18-cable duct

ArmorPVC

CabInner &

Xitu |~ ) Outer

Sheath

ThreeSymmetricalGrounds

Fig. 25 Aluminum armor cable ofField Site #2 showing A,B conducting

Fig. 23 shows measured I,, field site data for similar input rectifiertype loads, but different input cable geometry. Field Site #1 had an 18-conductor cable duct and Field Site #2 with Fig. 25 parallel 6 3-symmetric triangle bundled conductors, 3-symmetric ground Aluminumarmor 500 MCM cable used. Fig. 23 shows 1/2 the peak and rms Ismagnitudes for the Al armor cable. The 20 Amp peak-peak zerocrossing ripple, highlighted by Fig. 23a-23b dotted lines, represents themost symmetric two phase converter load operation of the cable duct asSection VI E explained. The 50 Apk level of Fig. 23a represents themost unsymmetrical two-phase operation of the cable duct.

Similarly, the "perfect 3-phase symmetry" of Fig. 25 is also alteredunder two-phase rectifier type loading. However, due to A, B, Ctriangular bundling and close conductor proximity, induced 20 Apk-pkI,s magnitude is similar in value to that for symmetric two-phaseoperation of Fig. 23a. In addition, any two phases conducting in Fig. 25will always have the same apparent symmetry under any 600 operatinginterval. This is a 3-triangular ground wire plane, with phase conductorson two sides of the triangle, with an overall aluminum armor ground.Thus, induced I s "humps" are equal during each 600 interval and resultsin the lumpy sinewave waveform of Fig. 23b.Table 7 compares I s measured in the PE tray ground wire at the two

field sites with similar common bus line-ups under similar loading.

Table 7 Measurement comparison of I,, in Fig. 4 cable tray PE bond wire forsame input rectifier load to MCC but different input cable geometry, 18conductor duct vs. parallel 11 6 - 500 MCM Al armor in tray

..... ......... ....... ... ...

... ....... ......... ...

LOADING Field Site #1 Field Site # 2 % ILs ComparisonIzs [Arms] I,s [ Arms] [#1/ #2] *100

MCC Load 1 61 27 225 %MCC Load2 42 18 233 %MCC Load3 32 16 200 %MCC Load 4 32 11 290 %

Average %0zs = 233%

Table 8 Simulation comparison of I, in Fig. 4 cable tray PE bond wire for1800 Arms input rectifier load to MCC but different input cable geometry,18 conductor duct vs. parallel 11 6 Al armor in tray and ground plane heights

GroundPlaneHeight [ft]

0 m Cm 1 m 2Ow 25m 3Qm 3 m 40m 5mt [$

(a) Simulated I, waveform at MCC Load 1

.~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~J..... :.::...

::

:....~ ~~~~~~~~~~ -:,, -.

:: ~ ~ ~~ ~ ~ ~ ~ ~~~~~~~ .,, .- -- ...

(b) Measured I, 27 Arms waveform at MCC Load 1

Fig. 26 Field Site # 2 aluminum armor cable in Al tray of Table 7 & 8

Induced I * is - 2.33x greater with the cable duct system than thealuminum armor in tray.

High magnitudes of Is can cause electro-chemical corrosion ofground grid joints. More importantly, uncontrolled I * current paths canflow through motor frames, shafts and bearings resulting in destructivepitting of the bearing race [15].

B. Simulated vs Measured I, at Field Site #1 & Field Site #2

Simulated Izs in Fig. 24 is close in waveshape quality to measured FieldSite #1 18-cable duct of Fig. 23a, with exception of peak currentmagnitude. Likewise, simulated vs. measured waveshapes are inagreement for Fig. 26. Field Site #2 aluminum armor cables inAluminum cable tray. Simulated peak magnitude variance in both casesis possibly due to parallel steel conduits below the tray area that raise theground plane height to Ift (from 8 ft) and serve as induced current pathsfor flux lines leaving the tray. Table 8 compares I * simulated in the trayground wire at the two field sites under similar loading. Simulated I,values are also 2.33x greater with the cable duct system than aluminumarmor in tray.

Field Site #1 Field Site # 2 % 1,, ComparisonIs [Apk] Is [Apk] [#1/ #2] *100

75

25

2.i

8foot 18 8 225 %Ifoot 27 11 245 %

Average %00, = 235 %

VIII. CONCLUSION

Common high-current cable geometries used in parallel circuits suchas Low Reactance (LR), Super Bundle (SB) and Triangular (TR)conductor phasing was reviewed for point and axial symmetry.

Evaluate cable geometry and conductor symmetry for linear or non-linear loads > 200 hp, since this hp range typically initiates parallelcircuit use. Asymmetrical arrangement of cable is the cause of zero-sequence current generation. Unbalanced magnetic fields due toconductor dis-symmetry start to produce noticeable induced zerosequence ground current (Izs) magnitudes at this hp and above.

This paper investigated a specific field site phenomenon in which alarge Izs current was found to circulate in the plant ground grid, eventhough the main supply transformer was ungrounded. Lack of conductorphase symmetry of the drive six-circuit input cabling in an aluminumtray resulted in an unbalance in mutual impedance coupling betweenparallel circuit conductors, so that a net external mutual magnetic fluxcoupling to the grounded cable tray existed. Induced (Izs) circulated inthe tray ends and PE copper grounds to building steel to form acirculating ground loop current, such as a single turn transformer loadwith the plant ground completing the secondary loop. Simulation with agrounded WYE source shows I * returns through the transformer neutral.

Magnetically induced ground current is influenced by both cablegeometry and operating characteristics of non-linear power converters.Zero sequence current is the vector sum of the zero sequence currentgenerated due to dis-symmetry in its own circuit geometry and the zerosequence currents induced from other parallel circuits. Zero sequencecurrents varied greatly according to cable phasing arrangement. Eachconductor phasing arrangement and overall cable geometry has aspecific IZS Iphase load ratio with a linear load. IZS Iphase load ratio isfurther increased under non-linear converter operation. Izs magnitude iscx load current so conductor symmetry at high current becomes critical.Parallel Two-Circuits laid in cable duct tray can have zero induced

ground current with linear loads by insuring mutual impedance' sbetween conductors are balanced by selecting point or axial symmetricconductor phasing arrangements of the six-conductors in the two parallelcircuits. Planar arrangement is acceptable if it is axial symmetric.

Parallel Four-Circuits laid in a cable duct tray will have Izscirculating current that cannot be reduced to zero with linear loads, evenwith point or axial symmetry conductor configurations. While each 2-circuit set can be arranged for zero Izs, there is always an unbalancedmutual coupling between the four parallel circuits and a zero sequencecurrent induced from the opposite set appears. (Izs / Iload) ratio of 4 % to14 % can be expected depending on conductor arrangement in the tray.Izs reduction requires conductor transposition in a four-circuit system.Parallel Five or Six -Circuits laid in a cable duct tray will always have

Izs circulating current that cannot be reduced to zero with linear loads.The parallel six-circuit bus duct is not perfect symmetry so inducedmagnetic flux lines spread out and into the high conductivity aluminumtray. A low voltage is induced along the tray, and since tray ends are PEwire bonded to earth ground, the resulting induced current can be large.Parallel Six-Circuits laid in a cable duct tray was investigated in

detail. Separate FEA analysis of the cable duct tray showed lowest eddycurrent loss induced in the tray was for an Inverse LR phased acacac,bbbbbb, cacaca arrangement, whether the tray was aluminum or steel.The FEA program generated an Impedance Coupling Matrix modelbased this geometry that represented the unbalanced mutual couplingbetween the parallel circuits, which was further embedded into a circuitsimulator program to determine induced ground current magnitude andwaveform in the PE tray end bond wire. Linear load simulation showeda 0.7 % Izs / Iload ratio can be expected with this geometry. Flux plotsshowed flux lines extend 4 feet beyond the six-circuit aluminum tray fora high current rectifier load, inducing measured Izs into adjacent steelconduits. Measured Izs /Iload ratio of 3.4 % to 4.8% occurred under non-

linear power converter operation. The paper successfully simulated theinduced zero sequence ground current phenomenon with powerconverter loads by showing similar waveform characteristics. Itadditionally showed further unsymmetrical duct operation, sincerectifier line currents conduct only two phases vs. three at a time forlinear loads. Exact peak current magnitude simulation was not asaccurate due to complexity of an industrial plant ground system.Parallel Six-Aluminum Armor Cables laid in an aluminum cable tray

was also investigated for similar loading at another site. Measured Izs /Iload ratio of 1.82% occurred with non-linear power converter operation.Lower Izs results since the cables are triangular bundled with 3-symmetric grounds. However, even the "perfectly symmetrical" cablegeometry is also unsymmetrical under two-phase converter operation.Flux plots also show unbalanced cable-cable circuit coupling still existssince the armor is non-magnetic. Measured and simulated results bothshow a 225 % reduction Izs magnitude compared to cable bus ductsystem and an external magnetic field reduction outside the tray of 12x.In contrast, no low frequency electromagnetic induced Izs on inverteroutput cables was found with Al armor since all three-phase currents aresinusoids and conduct simultaneously. Electro-static capacitive couplingswitching noise currents need a low resistance armor back to the VFD.

Parallel Galvanized Steel Armor Cables with triangular bundling, 3-symmetric grounds and outer PVC jacket warrants future investigationfor converter input wire use in cable tray. Cable-to-cable mutualinductive coupling is minimized making each symmetric circuit operatemore independently. However, there is still some cable dis-symmetryunder two-phase converter operation External magnetic field reductionoutside the tray is JOOx lower than the aluminum duct system. Inverterinduced electro-static switching noise can return back to the VFD withother filters, so that the steel armor benefits on the input side is possible.

Acknowledgements:Authors wish to thank B. Van Lieshout, B. Davis, K. Philips & J. Simons forpaper support

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