nafems thermal performance evaluation of air circuit ...

NAFEMS World Congress 2015 inc. the 2nd International SPDM Conference | San Diego, CA, 21-24 June 2015

NAFEMS THERMAL PERFORMANCE EVALUATION OF AIR CIRCUIT BREAKER (ACB) USING COUPLED ELECTRIC-THERMAL ANALYSIS Vikrant Deshmukh, Arnab Guha ,Sonil Singh, Subhash NN (EATON India Engineering Centre, INDIA); Ramdev Kanapady, Ph.D, ASME Fellow (EATON, USA). Abstract Air circuit breaker (ACB) protects downstream electrical systems and personnel from fault conditions due to abnormal current. The abnormal current arises due to overload faults, short circuit faults and ground faults. The size of ACB has been decreasing in accordance with voice of customers, which resulted in higher power density and internal heat generation which causes higher component temperature. The heat sources of ACB are the current carrying conductors and their joints. Therefore, proper knowledge of thermal characteristics of ACB current conducting path is essential for the accurate thermal design to ensure safe and reliable operation. This paper describes a new iterative sequentially coupled electricthermal finite element modeling methodology to predict the temperature rise in ACB. The proposed approach 1) captures the effect of resistance rise with temperature, 2) accurately accounts for electrical contact resistances which are modeled as contact volumes validated by measured values, 3) incorporates accurate heat transfer coefficients required at surfaces of inside the circuit breaker current path, heat dissipation from the plastics and mounting frame which are obtained from a separate CFD model that account for both convective cooling and radiative heat transfer and 4) inclusion of thermal contact resistance at various contact surfaces to accurately model the heat barriers to the heat flow along the path of the current path. The percentage error in predicting the temperature rise (under prediction) can be as high as 28% if the effect of resistance rise with temperature is ignored. The proposed approach is validated with temperature rise test results and the overall prediction is within ±5% error.

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1.

Introduction

Circuit breakers protect the electrical equipment, electrical wiring/cabling system and personnel from fault conditions due to excessive or abnormal current. The abnormal current arises due to overload faults, short circuit faults and ground faults. Circuit breakers are manufactured in wide variety of forms depending on their functionality or application, from small devices for a single household appliance up to large puffer circuit breakers for high voltage power transmission for an entire city. Low-voltage circuit breakers (LVCB) are common in domestic, commercial and industrial applications in the range up to 1000 VAC. LVCB can be a miniature circuit breaker (MCB) with rated currents less than 100A are used in domestic, commercial and light industrial applications. Molded case circuit breaker (MCCB) with rated currents less than 2500A used in low voltage switchboards, motor control center and panel boards. Low voltage power circuit breakers (LVPCB) with rated currents in the range 800 to 6300A are used in multi-tiers low voltage switchboards or switchgear cabinets. Insulated case circuit breaker (ICCB) in low voltage switchboard is same as LVPCB used in motor control centers and some transfer switches. LVPCB are also called as air circuit breakers (ACB). General discussions of the low-voltage circuit breaker and the principle of working can be found in the literature for instance in the book edited by Flurscheim [1]. Schematic description of single pole of an ACB is depicted in Figure 1 showing the different elements of current path.

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Figure 1:

Different elements of current path in air circuit breaker.

The continuous current rating of a circuit breaker is the number of amperes that the device can carry continuously without the temperature of any insulation of component becoming greater than its rated temperature. The maximum permissible temperature rise at rated current is defined and limited by international standards such as IEEE C37.13 and IEC 60947-2 [2] for circuit breakers. According to IEC 60947-2 temperature rise shall not exceed the values given in Table 1. Temperature rise in various components of the ACB are due to Joule heat generated by the specific resistance of each individual part and the contact resistance at the contacts, junctions, connections and bolt joints. Referring to Figure 1, the current path include line-side terminal, line-side junction, line-side conductor, main contacts, movable conductor, flexible conductor, load-side conductor, load-side junction and load-side terminal, in that order. Along this path it is important to identify heat sources, heat sinks and heat barriers to accurately predict temperature rise of each individual part so that hot spots can be identified. In addition prediction capability will also help in overcoming correct topology identification of circuit-breakers to be installed inside switchgear assembly for given maximum continuous current and for specific switchgear operating/boundary conditions. This paper focus on developing three dimensional (3D) finite elements based coupled electric-thermal model using Maxwell [3] and Ansys [4]. This is two-way iterative sequentially coupled thermal-electric model accurately predict the temperature rise at rated current. Results presented for low voltage power circuit breaker with an operating voltage range between 250-600V and rated current of 4000A. The organization of this paper is as follows. A brief introduction and motivation was described in Section 1. In Section 2 previous work on thermal modeling of low-voltage circuit breaker is given. Section 3 presents the governing equations pertaining to electrical and thermal physics with boundary conditions followed by solution strategy for the proposed methodology. Illustrative example and validation are presented in Section 4. Concluding remarks are provided in Section 5. Table 1: Temperature rise limits for terminals and accessible parts [2].

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Description of Part * Terminals for external connections

Temperature-rise limits ** K 80

Manual operating means Metallic 25 Non-metallic 35 Manual operating means Metallic 40 Non-metallic 50 Manual operating means Metallic 50 Non-metallic 60 * No value is specified for parts other than those listed but no damage should be caused to adjacent parts of insulating materials ** The temperature rise limits specified are not intended to apply to new sample, but are those applicable to the temperature-rise verifications during the appropriate test sequences specified in clauses of IEC 60947-2

2.

Review of Previous Work

It is common practice in the ACB industry to employ thermal network theory techniques such as described in Cherukuri [5] to predict the temperature rise of the conductors and terminals of the circuit breakers. However, this technique limits the prediction capability during the design phase when the conductors in the current paths have complex geometry and shapes. Plesca [6] presented a thermal model based on the finite element method (FEA) for an ACB where measured values of electrical resistances from test data was used to model the heat sources thereby assuming the current densities are uniform in the current carrying conductor cross sections. This technique not only limits the designers to optimize the weight of copper conductors based on the current density distributions (current density crowding etc.) but also temperature effects on the resistance are assumed to be negligible, under-predicting the temperature rise. The percentage error can be as high as 28% for the temperature rise requirements specified in the standards [2]. Frei et al [7] reported FEA thermal model for MCCB employing direct electric-thermal multi-physics coupling to predict the temperature rise of components. Plastic molded parts are included in the model along with the conductors. Thermal and electric contact resistances were calculated analytically and included in the model. The convective heat dissipation is considered only from the molded plastic casing to the surroundings and convective heat dissipation from current path components to inner air cavity is assumed to be negligible based on the work of Barcikowski [8] and Barcikowski et al. [9] for relatively similar size of MCCB. In this paper, focus is ACB (LVPCB) where the inner cavity size is relatively large compared smaller size MCCB, heat

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dissipation by natural convection in the inner air cavity is included in the model with heat transfer coefficients calculated from a separate CFD model. Dilawer et al. [10] presented a direct multi-physics electricthermal analysis for predicting the thermal behavior of ACB. Electrical contact resistances were modeled as volumes and experimentally measured data was employed. Effects of thermal resistance were neglected and heat transfer coefficients values for convective cooling were adjusted to correlate the predicted temperature results with test results. This paper describes a two-way iterative sequentially coupled thermalelectric model to accurately predict the temperature rise of the current path components. Direct current (DC) conduction electrical analysis was performed using MAXWELL [3]. The heat sources of Ohmic losses are sequentially used as the input to steady state thermal model in ANSYS [4]. The novel features of the proposed model are 1) capturing the effect of resistance rise with temperature, 2) accurately accounted electrical contact resistances which are modeled as contact volumes validated by measured values, 3) incorporate accurate heat transfer coefficients required for surface heat dissipation inside the ACB current path cavity and heat dissipation from the plastics and mounting frame that are obtained from a separate CFD model that account for both convective cooling and radiative heat transfer and 4) inclusion of thermal contact resistance at various contact surfaces to accurately model the heat barriers to the heat flow along the path of the current path. The proposed approach is validated with temperature rise test results.

Figure 2:

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Sequentially Coupled Electric-Thermal Model Solution Methodology for Air Circuit Breakers.


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3.

Sequentially Coupled Electric-Thermal Model for Air Circuit Breakers

It is well known that the temperature distribution in the conductors of circuit breaker under continuous current is a very complex phenomenon, which involves the coupled interactions between thermal and electromagnetic processes. An applied current flowing in electrically conducting path causes Ohmic losses which cause the temperature rise. This generates heat which is transported by conduction along the conductor path and also dissipated by convection and radiation from the surfaces of conductors and parts connected to it. This rise in temperature additionally causes the Ohmic losses to increase due to increase resistivity of the copper. This coupled electricthermal physical phenomena in the proposed model is described in Figure 2. In addition, main contacts and contacts at the junctions (refer Figure 1) provides additional Ohmic losses due to contact resistance and barrier to heat path by thermal contact resistance. The governing equations and boundary conditions related these physical phenomena are described in this section. 3.1 Electrical Model and boundary conditions In this section differential equations and associated boundary conditions governing the electrical model are presented. Coupling variable, temperature, to the thermal model is also introduced. The skin effects at 60Hz were expected to be insignificant hence steady state analysis is considered. Electric current density (J) is related to electric field (E) by Maxwell’s equations = ( ) = −∇∅

(1) (2)

where is electric conductivity, ∅ is scalar electric potential. This equation is analogous to Ohm’s law, where E is analogous to voltage, J is analogous to current and electrical conductivity is the inverse of resistivity. The Maxwell’s equation states that electric field in a material with non-zero conductivity produce an electric current. The governing equation for steady state electrical conduction is given by ∇∙ =0

(3)

∇ ∙ ( ( )∇ ∅) = 0

(4)

subjected to boundary conditions ∅ = 0 at the ground surface and ∅ = . = at the terminal surface. In the electric FE model, current

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source values were imposed on the free end of line-side bus bars and ground surfaces are were imposed on the free end of load-side bus bars. Since electric conductivity significantly varies with temperature, can be estimated through a linear variation of resistivity as ( )=

( )

=

(5)

(

Where Tref is reference temperature for ρo and α is the temperature coefficient of copper (0.00386 per oC). The temperature variable T provides the coupling variable from the thermal model which is described in the Section 3.3. 3.2 Electrical Contact Resistance The contact resistance is one of the critical heat sources along with the conductor resistance. In a ACB contact resistances come into play at two contacts one at the main contact and another at the conductor and terminal junctions of both line and load side of the current path. The contact resistance between two any surfaces 1-2 can be expressed as ∙

!

= ℎ#$%&'#& (

∙

)

= ℎ#$%&'#& (

(

−

()

(6a)

−

)

(6b)

Where n is outward normal at the contact surfaces, Ji is the current density at the surface i, Vi is the voltage at the surface i between the surface in contact, hcontact is the coefficient which depends on the contact pressure, material hardness, electrical conductivities of the material in contact and surface roughness. Just like the resistivity of the conductors, all these parameters are function temperature. One could employ coefficients hcontact, in Eqs. 6 as described in [6, 11]. In this study we have employed experimentally measured contact resistances at room temperature to validate the proposed coupled electrical-thermal model. 3.3 Thermal Model and boundary conditions In this section differential equations and associated boundary conditions governing the thermal model are presented. Coupling variable, current densities, to the electrical model are also introduced. For a steady state conduction problem the differential equations governing the thermal response of the system under thermal loads and associated boundary conditions are −* +

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,-

,. -

+

,-

,0 -

+

,-

,1 -

2 = 3(4)

(7)


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subjected to following boundary conditions Essential BCs:

=

,

(8a)

'

Neumann BCs: −* + 2 = 56 ,%

Convection BCs: −ℎ( −

7)

= 56

(8b) (8c)

where qb is heat flux (W/m2), K is thermal conductivity (W/mK) and n is the outward unit normal to the surface and Q(x) is Ohmic losses given by 3(8) =

|:(.)|-

(9)

;( )

where J(x) is the current density which provides the coupling variable from the electrical model. 3.4 Thermal Boundary Conditions Heat transfer due to convection (qconv) from the current path conductors to the surrounding is modeled as convective boundary and is defined by in the thermal model by heat transfer coefficient (HTC), surface area of components (A) and temperature difference given by (

?@AB'#C

−

'DE7C%& )

(10)

Where h(x) is the HTC which varies along the surfaces of conductor, A is the surface area of the conductor and Tsurface is the surface temperature of the conductors and Tambient is the ambient temperature. In the next section the methodology to calculate the HTC values from a circuit breaker CFD model is described and the calculated values are presented in Section 4. 3.5 CFD Model for Heat Transfer Coefficients The HTC defined in Eq. 10 is obtained from CFD model with steady state conjugate heat transfer model with radiation. The heat sources are included as Ohmic losses in the current path conductors with the assumption that the current densities are uniformly distributed across the cross section of the conductors (no coupling with electric model). External air domain extends beyond the ACB frame. Boundary conditions for the external domain is set as atmospheric pressure (P = Patm) and reverse flow temperature equal to atmosphere (T = Tatm). Air has been modelled as incompressible ideal gas to allow change in density only due to temperature.

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Radiation has been modelled using S2S model. In this method view factor is calculated before the simulation begins and extent of radiative heat exchange is calculated with respect to view factor. In S2S model, view factor for finite surfaces i with respect to j is given as F7G =

HI

JH JH I

K

LMN OI LMN OK PA -

Q7G RS7 RSG

(11)

Where, r is the distance between the surface centroids, θ is the angle between lines joining the surface centroids and their respective normal and δ is equal to 1 if surfaces are visible to each other and 0 otherwise. Turbulence has been modelled using standard K-epsilon RANS model. Standard K-epsilon is considered as this model is being widely used industrial applications. Standard K-epsilon turbulence (Launder & Sharma) models turbulent viscosity using equations for Turbulence kinetic energy (k) and dissipation (Є) as given below. T

TU

T

TU

(VW) +

(VЄ) +

T

T4I

T

T4I

(VWX7 ) =

(VЄX7 ) =

T

T4K

T

T4K

[

T]

Y+Z + ; I 2 T4 ^ + _] + _E - VЄ – cd + e] \

[

TЄ

Y+Z + ;I 2 T4 ^ + f Є

K

K

Є Є ] (_]

(12a)

+ fgЄ _E ) − f(Є V eЄ

Є]

+

(12b)

The HTC values are calculated using ℎ(4) = 5/( i'jj − 'DE7C%& ), where q is the heat flux and Twall is the surface temperature of the conductor. Advantage of using CFD thermal model is elimination of need to approximate HTC at conductor surfaces as it includes buoyancy effects. But due to higher computational effort required, CFD has been limited to one time analysis to calculate HTC distribution along the surfaces of the conductors. Since radiation is included in the CFD model, this unique approach alleviates need of modelling radiation in conduction thermal model. The HTC values from CFD not only improve fidelity but also reduce time required in conduction thermal model. The HTC employed in the model are presented in Section 4. 3.6 Thermal Contact Resistance Thermal contact resistance plays equally important role as electrical contact resistance in determining temperature rise of the components. Depending upon the medium between the contacting surfaces may contribute or restrict the heat transfer at the junction. There is no temperature drop at the interface of bolted joints. Hence it is assumed that there is perfect thermal contact conductance (TCC) between these joints. But at the contacts between moving arms-spacers (main

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contacts) and fingers-conductors (auxiliary contacts), the heat flow across the contact surface is defined in ANSYS using < = ff

&'AkC&

−

(13)

#$%&'#&

where TCC is the thermal contact conductance, Ttarget and Tcontact are the temperature of corresponding nodes at the contact surfaces. Negus [12] presented a model for TCC of Eq. 13 as ff = 1.25W?

D ?

o r.st

+ 2

(14)

pq

where P is apparent pressure on interface, Hc is hardness of softer material, ks is effective thermal conductivity, m is effective absolute (] ] surface slope and s is effective surface roughness given as W? = ] u ]- , u

-

v = wv + v( x = wx + x( For most practical applications o involving both similar and dissimilar materials, 10 y ≤ p ≤ 2 ∗ 10 ( , (

(,

(

(.

q

x7 = 0.125 (v7 ∗ 10y )r.|r( for 0.216 ≤ v7 ≤ 9.6 Zx can be employed. Using Eq.14 TCC is calculated and incorporated in ANSYS thermal model. 4.

Simulation Results

The Figure 3 shows the current path components of the ACB that is being investigated in this study. The 3D CAD model created in Pro-E was imported to Maxwell. DC conduction analysis was performed in Maxwell and the heat load data was imported to the ANSYS thermal module.

Figure 3:

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Terminal Resistance measurement points.


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HTC values obtained from CFD analysis was used as input in the ANSYS thermal model. The HTC coefficients obtained from CFD model account for heat loss due to both convection and radiation. The Figure 4 shows the HTC distribution along the surfaces of conductor which is employed for the validation of the model. In the ACB configuration, the line-side components is vertically above the load-side components, hence CFD analysis correctly predicts that the HTC values of line-side components are lower from the load-side components. The maximum HTC value is 12 W/m2K. From the Figure 4 it is clear that there is heat dissipation from the current path components to inner cavity and needs to be included in the model to accurately predict the temperature rise of the ACB components.

Figure 4:

Heat Transfer Coefficient Distribution for the Current Carrying Components of ACB.

The contact resistances need at surfaces of the moving arm and the line conductor, and at the primary disconnects which connect conductor extensions and adaptors both at line and load sides. The electrical contact resistances were derived from the resistances of lab measurements between points A-B, C-D and E-F shown in Figure 4 and listed in Table 2. In the electrical model without thermal model, the contact parameters were varied so that total electrical resistances of 3 measurements matched with lab test results. This resulted in main terminal contact resistance of 2.5 µΩ and junction contact resistance of 0.5 µΩ between primary disconnects and extensions and adaptors. Table 2: Test and simulation terminal resistances.

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Measurement points

Resistance (µΩ)

A-B

14

C-D

10.3

E-F

7

Table 3: Material properties for thermal model.

Material

Thermal Conductivity (W/mK)

Copper

390

Plastic

0.6

Steel

16

In ANSYS thermal model material properties and TCC were assigned to the ACB components. The material properties used for thermal analysis is shown in Table 3. For electrical analysis a reference electrical conductivity of 5.8×107 S/m is used for copper. Using Eq. 14 resulted in TCC of 7000 W/m2 0C employed for main contacts and junction contacts.

Figure 5:

(a) Ohmic loss and (b) temperature distribution.

DC Current excitation of 4000A was applied at the bus bar free end and the coupled analysis was performed analysis for single pole of the ACB. Proximity effects among poles were assumed negligible. Analysis was

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performed for several iterations. After five cycles of iteration, there was no significant change in temperature values. The Figure 5a shows current density distribution pattern and the Figure 5b shows the final temperature distribution. The percentage error in temperature rise prediction and lab test data are shown in the Figure 6 for various components along the current path. In Figure 6, sequentially coupled results corresponds to current proposed approach, one way coupled results corresponds to no increase of resistivity with temperature but the Ohmic losses are imported from Maxwell at room temperature and no coupling results corresponds to only thermal model with Ohmic losses which are uniformly distributed in the current carrying conductors. For sequentially coupled approach there is a good correlation between experiments and FE results are within ±5% error. The results from the study bring out the following inferences. 1. If the of effect increase in resistance with temperature rise is not included in the model, the percentage error of under predicting the temperature rise (under prediction) can be as high as 26% and 28% for models with one way coupling and without coupling with electric model respectively. The proposed two-way iterative sequential coupling of electric and thermal model effectively captured effect of increase in resistance with temperature rise. 2. The internal heat generation and resulting operating temperature of current path critically depends on the electrical contact resistance. Hence for temperature rise prediction model to be effective a validated contact resistance model is needed. 3. The percentage error between the predicted and the test results are significant for conductors inside the ACB if thermal contact resistance is not included in the model. The temperature rise will be lower if the TCC is high. The TCC depends on various control factors such as force on contact interface, hardness of materials, radius of contact spot, surface roughness of materials, thermal conductivity of materials in contact and void area. 4. HTC values from CFD increases the thermal prediction performance significantly. If time and resource permits, then sequentially coupled electric-CFD model needs to be employed as this model will allow to explore further cooling options related to internal chamber of ACB to bring down the temperature rise.

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Figure 6:

5.

Percentage error of temperature rise results with lab test results for components along the current path.

Conclusions

A new iterative sequentially coupled electric-thermal finite element modeling methodology to accurately predict the temperature rise for ACB is presented. For the first time, incorporating all the following key features are proposed for ACB. The key features are 1) captured the effect of resistance rise with temperature, 2) accurately model of electrical contact resistances which are modeled as contact volumes validated by measured values, 3) incorporated accurate heat transfer coefficients required for surface heat dissipation inside the circuit breaker current path cavity and heat dissipation from the plastics and mounting frame are obtained from a separate CFD model that account for both convective cooling and radiative heat transfer and 4) inclusion of thermal contact resistance at various contact surfaces to accurately model the heat barriers to the heat flow along the path of the current path. The percentage error in predicting the temperature rise (under prediction) can be as high as 28% if the effect of resistance rise with temperature is ignored. The proposed approach is validated with temperature rise test results and the overall prediction is within ±5% error. The developed model will allow us perform comprehensive study

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on thermal performance for design changes such as but not limited to 1) the effect of various control factors such as line bus bar size, number and arrangement and 2) the effects of various noise and control factors pertaining to thermal contact conductance and electrical contact resistance. 6.

Acknowledgements

Authors would like to thank entire ACB design team at Eaton, especially Paul Rakus and Dr. Michael Slepian for their ACB product technical inputs to this project.

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7.

References

1.

Flurscheim, Charles (1985), Power Circuit Breaker Theory and Design, Peter Peregrinus Ltd., London, 2nd Edition.

2.

IEC 60947-2 Low-voltage Switchgear and Controlgear – Part 2: Circuit Breakers Edition 4.2 2013-01.

3.

ANSYS Maxwell 3D v16: user Guide.

4.

ANSYS: Inc. Theory Reference [online].

5.

Cherukuri, Krishna Swamay (2007), Thermal Network Theory for Switchgear under Continuous Current, MS Thesis, National Institute of Technology, Rourkela.

6.

Plesca, Adrian (2012), Thermal analysis of the current path from circuit breakers using finite element method: World academy of science, Engineering and Technology, Vol: 72, 12-29, 2012.

7.

Frei, Peter and Weichert, Hans (2004), Advanced thermal simulation of a circuit breaker, Electrical Contacts: Proceedings of the 50th IEEE Holm Conference on Electrical Contacts and the 22nd International Conference on Electrical Contacts.

8.

Barcikowski, Frank (2003), Numerische Berechnungen zur Wärme und Antriebsauslegung von Schaltgeräten, Ph.D. dissertation, Technische Universität Braunschweig, Göttingen: Cuvillier.

9.

Barcikowski, Frank and Lindmayer, Manfred (2000), Simulations of the heat balance in low-voltage switchgear, in Proc. 20th Int. Conf. on Electrical Contacts, Stockholm.


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10. Dilawer, Syed Ibrahim, Junaidi, Md. Abdul Raheem, Samad, Mohd Abdul, and Mohinoddin, Mohd. (2013), Steady state thermal analysis and design of air circuit breaker: International journal of engineering & technology, Vol. 2, Issue 11. 11. Holmes, Ragnar, Electric Contacts: Theory and Applications, Springer-Verlag, 4th Edition. 12. Negus, K.J. and Yovanovich, M.M. (1988), Co relation of the gap conductance integral for conforming rough surfaces: Journal of Thermo Physics, Vol. 2, No 3. 13. http://www.mecheng.osu.edu/documentation/Fluent14.5/145/wb_si m.pdf, 22 July 2014.

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