# Investigation of Thermal Effects in Different Lightweight Constructions for Vehicular Wireless Power Transfer Modules

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## Abstract

**:**

## 1. Introduction

## 2. Exemplary Wireless Power Transfer System

## 3. Description of Proposed CPM Concepts

#### 3.1. Sandwich Concept

#### 3.2. Space-Frame Concept

## 4. Thermal Model

#### 4.1. Calculation of Power Losses

#### 4.1.1. Coil Losses

#### 4.1.2. Core Losses

#### 4.1.3. Eddy Current Losses

#### 4.2. Thermal Simulation Model

^{2}K according to [42]. Second, forced convection is applied adjusting heat transfer coefficient $\alpha $ to 25 W/m

^{2}K taken from [43]. Analogous to literature, heat transfer by heat radiation is neglected [19,44]. The used material parameters of the thermal model are summarized in Table 4.

## 5. Experimental Setups

#### 5.1. Component Level Testing

#### 5.2. System Level Testing

## 6. Results

#### 6.1. Sandwich Concept

**Component level testing:**Figure 12 summarizes the results of the thermal investigation of the sandwich concept, taking into account the electric requirements of component level testing (${I}_{\mathrm{CPM}}=33\phantom{\rule{0.166667em}{0ex}}\mathrm{A}$) and natural convection. The thermal limit ${T}_{\mathrm{Max}}$ of this concept is reached within a time of 120 min shown in Figure 12b. In this respect, the ferrite temperature ${T}_{\mathrm{fer}}=130.5\xb0\mathrm{C}$ exceeds the temperature limit of the used epoxy CW 2250-1 under natural convection. Aside from the coil, the 3F36 ferrites contribute a significant thermal load, wich is shown in Figure 12a. This can be reduced using ferrites more appropriate to the chosen frequency. As expected, the lower thermal conductivity of the foam core impedes the heat transfer to the bottom side of the prototype.

^{2}K is considered to be valid. To evaluate the thermal feasibility of the sandwich concept, a FE analysis with forced convection is applied adjusting the heat transfer coefficient $\alpha $ to 25 W/m

^{2}K. The active cooling results in a significant reduction of the steady-state temperatures shown in Figure 12d. The temperatures of coil and ferrites are not critical regarding the thermal limit of the used thermal conductive resin. Thus, the thermal feasibility of the proposed sandwich concept using active air cooling is considered to be valid.

**System level testing:**Figure 13 summarizes the results of the thermal analysis of the sandwich concept, taking into account natural convection and the electric requirements of system level testing. A slightly higher battery voltage of this setup results in lower CPM current (${I}_{\mathrm{CPM}}=29\phantom{\rule{0.166667em}{0ex}}\mathrm{A}\phantom{\rule{4.pt}{0ex}}\mathrm{RMS}$). The results of the coil temperature measurement are presented in Figure 13c. The sensor positions and the results of FE-analysis are shown in Figure 13a. As predicted, the substitute modeling of the coil in the FE model makes the assessment of the temperature gradient across the coil windings impossible. As a result, the temperature rise of the outer windings is mapped accurately, whereas the discrepancy increases towards inner windings. Moreover, compared to the component level testing, this effect might be amplified by the additional losses caused by the magnetic field of the GPM. However, the average accuracy of the proposed FE model is approximately 5 K in the critical coil temperature. The results of the FE analysis and experimental validation of the shielding temperatures are shown in Figure 13d. The simulation result of the shielding temperature is calculated from the mean of the shielding surface. The experimental validation shows a significant temperature gradient towards the center of the shielding surface temperature. Due to uniform distribution of the power losses over the shielding surface in the FE model, this is not reproduced. In this respect, a fully coupled electromagnetic–termal simulation with a highly decrete shielding model might be more accurate. However, the FE model reproduces the measured temperatures of the shielding edge well, which is crucial for thermal management. In this respect, the simulation results in Figure 13a and the thermographic image of Figure 13b show a good consistency. Taking into account the testing specifications of SAE J2954 [5] and the system specifications of Table 1, the DC-DC efficiency of the sandwich concept is measured over the measured period according to Figure 13c at an input power of 11 kW between 88.7% and 91.7%.

#### 6.2. Space-Frame Concept

**Component level testing:**Figure 14 summarizes the results of the thermal analysis of the space–frame concept. Taking into account the electric requirements of component level testing (${I}_{CPM}=33\phantom{\rule{0.166667em}{0ex}}\mathrm{A}\phantom{\rule{4.pt}{0ex}}\mathrm{RMS}$) and natural convection, we measured the temperature rise within a time of 70 min. Compared to the sandwich concept, the improved heat transfer to the bottom and top side combined with the bigger heat sinks enables a significant heat reduction. Consequently, the temperature gradient from center to top and bottom side decreases, shown in Figure 14a. Thus, the maximum operating temperature of the thermal conductive resin is not reached within the measuring time.

**System level testing:**Figure 15 summarizes the results of the thermal analysis of the space–frame concept. Considering the electric requirements of system level testing (${I}_{\mathrm{CPM}}=29\mathrm{A}\phantom{\rule{4.pt}{0ex}}\mathrm{RMS}$) and natural convection, we measured the temperature rise within a time of 120 min. As postulated in the previous section, the cutouts at the corners of the shielding result in local hotspots shown in the thermographic image of Figure 15b. Therefore, the temporary temperature along the shielding edge is not uniform. Thus, the temperature measurement using fibre optical sensors is pointless. Besides, the FE model does not reproduce the temperature rise of the heat sink according to the component level testing. Therefore, we only measured the temperature rise of the coil shown in Figure 15c. Analogous to the sandwich concept, the temperature rise of the outer windings is mapped accurately, whereas the discrepancy increases towards inner windings. However, the proposed FE model is sufficiently accurate in the critical coil temperature. However, proof of thermal feasibility requires further experimental validation with a completely closed shielding. Thus, the shielding temperature is to be expected as uncritical, similar to the sandwich concept. Taking into account the testing specifications of SAE J2954 [5] and the system specifications of Table 1, the DC-DC efficiency of the sandwich concept is measured over the measured period according to Figure 13c at an input power of 11 kW between 87.8% and 90.1%.

## 7. Discussion

## 8. Conclusions

## Author Contributions

## Funding

## Conflicts of Interest

## Abbreviations

AC | Alternating current |

CPM | Car Pad Module |

DC | Direct current |

EVs | Electric vehicles |

FE | Finite elements |

FOM | Figure-of-Merit |

GPM | Ground Pad Module |

HF | High frequency |

HiL | Hardware-in-the-Loop |

MOSFET | Metal–oxide–semiconductor field-effect transistor |

PFC | Power factor correction |

RMS | Root Mean Square |

WPTS | Wireless power transfer systems |

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**Figure 2.**z-dimension of different WPTS constructions: (

**a**) conventional construction for WPTS with a transmission power of 3.6 kW given in [20], (

**b**) conventional construction for WPTS with a transmission power of 11 kW e.g. given in [23] and (

**c**) proposed concept for WPTS with a transmission power of 11 kW.

**Figure 7.**(

**a**) Analytically calculated magnetic field $H\left(x\right)$ of individual conductors and derived external magnetic field ${H}_{\mathrm{ext}}$ from the mean value of $H\left(x\right)$ of section A-A in (

**b**) the simplified model with solid regions of coil and ferrites taking advantage of axial symmetry.

**Figure 8.**Power loss densities: (

**a**) ${p}_{\mathrm{coil}}\left(T\right)$, (

**b**) ${p}_{\mathrm{core}}\left(T\right)$ and (

**c**) ${p}_{\mathrm{eddy}}\left(t\right)$.

**Figure 9.**Experimental setup of the thermal analysis and position of the fibre optical thermal sensors exemplary on sandwich concept.

**Figure 11.**Photograph of the experimental setup of the system level testing exemplary on the sandwich concept.

**Figure 12.**(

**a**) Results of the thermal FE analysis of sandwich concept on component level at $t=120\phantom{\rule{0.166667em}{0ex}}\mathrm{min}$, (

**b**) comparision of heating results of FE analysis and component level testing with natural convection, (

**c**) comparision of cool-down results of FE simulation and component level testing with natural convection and (

**d**) comparision of FE analysis with natural ($\alpha =10\mathrm{W}/{\mathrm{m}}^{2}\mathrm{K}$) and forced convection ($\alpha =25\mathrm{W}/{\mathrm{m}}^{2}\mathrm{K}$).

**Figure 13.**(

**a**) Results of the thermal FE analysis of sandwich concept on system level at $t=100\phantom{\rule{0.166667em}{0ex}}\mathrm{min}$, (

**b**) thermographic image at $t=100\phantom{\rule{0.166667em}{0ex}}\mathrm{min}$, (

**c**) heating results of FE analysis and system level testing of the coil and (

**d**) the shielding under natural convection ($\alpha =10\mathrm{W}/{\mathrm{m}}^{2}\mathrm{K}$).

**Figure 14.**(

**a**) Results of the thermal FE analysis of space–frame concept on component level at $t=70\phantom{\rule{0.166667em}{0ex}}\mathrm{min}$, (

**b**) comparision of heating results of FE analysis and component level testing with natural convection ($\alpha =10\mathrm{W}/{\mathrm{m}}^{2}\mathrm{K}$), (

**c**) FE analysis with forced convection ($\alpha =25\mathrm{W}/{\mathrm{m}}^{2}\mathrm{K}$).

**Figure 15.**(

**a**) Results of the thermal FE analysis of space–frame concept on system level at $t=120\phantom{\rule{0.166667em}{0ex}}\mathrm{min}$, (

**b**) thermographic image at $t=120\phantom{\rule{0.166667em}{0ex}}\mathrm{min}$, (

**c**) heating results of FE analysis and system level testing of the coil under natural convection ($\alpha =10\mathrm{W}/{\mathrm{m}}^{2}\mathrm{K}$).

Parameter | Variable | Value | Unit |
---|---|---|---|

Output Power | ${P}_{\mathrm{out}}$ | 11 | kW |

GPM DC-Link Voltage | ${U}_{\mathrm{GPM},\mathrm{DC}}$ | 0–540 | V |

CPM DC-Link Voltage | ${U}_{\mathrm{CPM},\mathrm{DC}}$ | 300–470 | V |

Battery Voltage | ${U}_{\mathrm{batt}}$ | 300–470 | V |

Air Gap | z | 100–210 | mm |

Misalignment longitudinal | $\Delta x$ | ±75 | mm |

Misalignment transversal | $\Delta y$ | ±150 | mm |

Transmission Frequency | ${f}_{0}$ | 81.38–90 | kHz |

Area of CPM coil $(x,y)$ | ${A}_{\mathrm{coil},\mathrm{max}}$ | 300 × 300 | mm^{2} |

Parameter | Variable | Value | Unit |
---|---|---|---|

Coil length | ${L}_{\mathrm{GPM}}$ | 600 | mm |

${L}_{\mathrm{CPM}}$ | 300 | mm | |

Coil width | ${W}_{\mathrm{GPM}}$ | 750 | mm |

${W}_{\mathrm{CPM}}$ | 300 | µm | |

Number of windings | ${N}_{\mathrm{GPM}}$ | 7 | |

${N}_{\mathrm{CPM}}$ | 15 | ||

Self-Inductance | ${L}_{\mathrm{GPM}}$ | 77 | µH |

${L}_{\mathrm{CPM}}$ | 112 | µH | |

Mutual Inductance | M | 7.4–18.6 | µH |

Magnetic Coupling | k | 0.08–0.2 | |

AC-Resistance | ${R}_{\mathrm{AC},\mathrm{GPM}}$ | <40 | m$\mathsf{\Omega}$ |

${R}_{\mathrm{AC},\mathrm{CPM}}$ | 100 | m$\mathsf{\Omega}$ | |

RMS-Current | ${I}_{\mathrm{GPM}}$ | 35–100 | A RMS |

${I}_{\mathrm{CPM}}$ | 23–36 | A RMS |

Name | Parameter | Value | Unit |
---|---|---|---|

Transmission frequency | f | 85 | kHz |

Peak coil current | $\widehat{I}$ | 33 | A |

Magnitude of external magnetic field | ${H}_{\mathrm{ext}}$ | 6.62 | kA ${\mathrm{m}}^{-1}$ |

Number of strands in the litz wire | N | 15 | |

Total length of litz wire | ${l}_{\mathrm{litz}}$ | 12.7 | m |

Outer diameter of the litz wire | ${d}_{\mathrm{litz}}$ | 4.5 | mm |

Specific resistance of litz wire material at 20 °C [40] | ${\rho}_{\mathrm{litz},20}$ | 0.018 | $\mathsf{\Omega}$ mm ${\mathrm{m}}^{-1}$ |

Specific temperature coefficent of litz wire material at 20 °C [40] | ${\alpha}_{\mathrm{litz},20}$ | 0.004 | ${\mathrm{K}}^{-1}$ |

Density $\mathit{\rho}$ $\left[\raisebox{1ex}{$\mathbf{g}$}\!\left/ \!\raisebox{-1ex}{${\mathbf{cm}}^{\mathbf{3}}$}\right.\right]$ | Heat Capacity ${\mathit{c}}_{\mathbf{p}}$ $\left[\raisebox{1ex}{$\mathbf{J}$}\!\left/ \!\raisebox{-1ex}{$\mathbf{kg}\phantom{\rule{0.166667em}{0ex}}\mathbf{K}$}\right.\right]$ | Thermal Expansion Coeffient ${\mathit{\alpha}}_{\mathbf{th}}$ $\left[{\mathbf{10}}^{-\mathbf{6}}\phantom{\rule{0.166667em}{0ex}}\raisebox{1ex}{$\mathbf{1}$}\!\left/ \!\raisebox{-1ex}{$\mathbf{K}$}\right.\right]$ | Thermal Conductivity $\mathit{\lambda}$ $\left[\raisebox{1ex}{$\mathbf{W}$}\!\left/ \!\raisebox{-1ex}{$\mathbf{m}\phantom{\rule{0.166667em}{0ex}}\mathbf{K}$}\right.\right]$ | |
---|---|---|---|---|

Copper [45] | 8.92 | 385 | 17 | 381 |

BMF8 [46] | 4.8 | 750 | 1 | 4.2 |

3F36 [46] | 4.8 | 750 | 1 | 4.2 |

IG-F 71 [30] | 0.075 | 200 | 38 | 0.03 |

CW 2250-1 [31] | 1.58 | 1800 | 65 | 0.76 |

PA [47] | 1.14 | 1700 | 80 | 0.23 |

Organic sheet | 1.35 | 1280 | 16 | 0.28 |

Silicone [35] | 3.2 | 710 | 200 | 6 |

Aluminium [48] | 2.66 | 900 | 25 | 117 |

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**MDPI and ACS Style**

Zimmer, S.; Helwig, M.; Lucas, P.; Winkler, A.; Modler, N.
Investigation of Thermal Effects in Different Lightweight Constructions for Vehicular Wireless Power Transfer Modules. *World Electr. Veh. J.* **2020**, *11*, 67.
https://doi.org/10.3390/wevj11040067

**AMA Style**

Zimmer S, Helwig M, Lucas P, Winkler A, Modler N.
Investigation of Thermal Effects in Different Lightweight Constructions for Vehicular Wireless Power Transfer Modules. *World Electric Vehicle Journal*. 2020; 11(4):67.
https://doi.org/10.3390/wevj11040067

**Chicago/Turabian Style**

Zimmer, Steve, Martin Helwig, Peter Lucas, Anja Winkler, and Niels Modler.
2020. "Investigation of Thermal Effects in Different Lightweight Constructions for Vehicular Wireless Power Transfer Modules" *World Electric Vehicle Journal* 11, no. 4: 67.
https://doi.org/10.3390/wevj11040067