Ball grid array-module with integrated shaped lens for WiGig applications in eyewear devices

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Citation for published item:

Bisognin, A., Cihangir, A., Luxey, C., Jacquemod, G., Pilard, R., Gianesello, F....Whittow, W. (2016). Ball grid array-module with integrated shaped lens for WiGig applications in eyewear devices. IEEE Transactions on Antennas and Propagation. 64 (3), 872-882

Further information on publisher's website:

10.1109/TAP.2016.2517667

Publisher's copyright statement:

This is the peer reviewed version of the following article: Bisognin, A., Cihangir, A., Luxey, C.,

Jacquemod, G., Pilard, R., Gianesello, F....Whittow, W. (2016). Ball grid array-module with integrated shaped lens for WiGig applications in eyewear devices. IEEE Transactions on Antennas and

Propagation. 64 (3), 872-882, which has been published in final form at

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Lens for WiGig Applications in Eyewear Devices

2

Aimeric Bisognin, Student Member, IEEE, Aykut Cihangir, Cyril Luxey, Senior Member, IEEE,

Gilles Jacquemod, Member, IEEE, Romain Pilard, Member, IEEE, Frédéric Gianesello, Member, IEEE,

Jorge R. Costa, Senior Member, IEEE, Carlos A. Fernandes, Senior Member, IEEE,

Eduardo B. Lima, Student Member, IEEE, Chinthana J. Panagamuwa, Member, IEEE,

and William G. Whittow, Senior Member, IEEE

3

4 5 6 7

Abstract—A ball grid array-module (BGA-module)

incorpo-8

rating a low-cost shaped dielectric lens is proposed for wireless

9

communications in the 60-GHz WiGig band between a smart

eye-10

wear, where it is integrated and facing a laptop or TV. The module,

11

which is codesigned with a 60-GHz transceiver, consists of two

12

separate identical antennas for transmitting (Tx) and receiving

13

(Rx). The in-plane separation of these elements is 6.9 mm both

14

being offset from the lens focus. This poses a challenge to the lens

15

design to ensure coincident beam pointing directions for Rx and

16

Tx. The shaped lens is further required to narrow the angular

17

coverage in the elevation plane and broaden it in the horizontal

18

plane. A 3-D-printed eyewear frame with an integrated lens and

19

a recess for proper BGA-module integration is fabricated in

ABS-20

plastic material. Measurements show a reflection coefficient below

21

−12 dB in the 57–66 GHz band. A maximum gain of 11 dBi is

22

obtained at 60 GHz, with 24◦ and 96◦ beamwidth at 5-dBi gain,

23

respectively, in the vertical and horizontal planes. The radiation

24

exposure is evaluated for a homogeneous SAM head phantom and

25

a heterogeneous visible human head. The simulated power

den-26

sity values for both models are found to be lower than the existing

27

standards.

28

Index Terms—60 GHz, antenna-in-package, eyewear, lens

29

antennas, plastic packaging, WiGig.

30

I. INTRODUCTION 31

W

ITH the never-ending improvement of the capabilities Q1 32

of wireless communication devices, the most critical 33

necessity has been to supply the user with higher and higher 34

data rates. This has led to both the improvement of the existing 35

Manuscript received November 16, 2014; revised November 05, 2015; accepted December 19, 2015. Date of publication XXXX XX, XXXX; date of current version XXXX XX, XXXX.

A. Bisognin is with EpOC, Université Nice Sophia Antipolis, Valbonne, France, and also with STMicroelectronics, Crolles, France.

A. Cihangir and G. Jacquemod are with EpOC, Université Nice Sophia Antipolis, Valbonne, France.

C. Luxey is with EpOC, Université Nice Sophia Antipolis, Valbonne, France, and also with Institut Universitaire de France (IUF), Paris, France (e-mail: cyril.luxey@unice.fr).

R. Pilard and F. Gianesello are with STMicroelectronics, Crolles, France. C. Fernandes and E. Lima are with Instituto de Telecomunicações, Instituto Superior Técnico, University of Lisbon, Lisbon, Portugal.

J. Costa is with DCTI, Instituto Universitário de Lisboa (ISCTE-IUL), and Instituto de Telecomunicações, Lisbon, Portugal.

C. Panagamuwa and W. Whittow are with the Wireless Communications Research Group, School of Electronic, Electrical and Systems Engineering, Loughborough University, Loughborough, LE11 3TU, U.K. (e-mail: w.g.whittow@lboro.ac.uk, C.J.Panagamuwa@lboro.ac.uk).

Color versions of one or more of the figures in this paper are available online at http://ieeexplore.ieee.org.

Digital Object Identifier 10.1109/TAP.2016.2517667

wireless communication standards as well as the launch of 36 new standards and new technologies. One of these standards, 37 the WiGig IEEE 802.11ad, is gaining more and more popu- 38 larity among industries because the unlicensed frequency band 39 around 60 GHz offers a broad bandwidth to achieve multigiga- 40 bits speeds (up to 7 Gbit/s). The low interference level favored 41 by the very high wall penetration loss and by the high oxygen 42 absorption in this band for moderate distances makes this stan- 43 dard a good candidate for line-of-sight (LoS) in-room wireless 44 personal area communications (WPAN). Possible applications 45 include the wireless connection of a personal computer (PC) 46 with its peripheral devices (monitor, keyboard, etc.), as well 47 as ultra-high-definition video/audio transfer from a camera to 48 a TV or projector, eliminating the need for cables. Typically, 49 for a LoS 2-meter communication in this band, an antenna gain 50 of approximately 4 dBi is needed (at both sides of the link) 51 considering today’s transceiver performances (10 dBm power 52 at the antenna port,−54 dBm Rx sensitivity and OFDM 16- 53 QAM modulation). If this distance is increased to around 8 m, 54 the gain should be approximately 10 dBi. 55 In parallel, smart eyewear devices are gaining popularity as 56 wireless communicating objects with some products already 57 released in the market and some other being prepared for the 58 near future [1]–[5]. In general, those devices incorporate a small 59 optical lens-reflector screen, a camera, a microphone/speaker 60 pair, and a touchpad. They are generally connected to a periph- 61 eral (smartphone or set-top box) through Bluetooth or WLAN 62 standards at 2.4 GHz. Our recent work considered eyewear 63 devices as a possible candidate to replace smartphones in the 64 near future and we successfully demonstrated high potential for 65

LTE communications [6]. 66

In this study, a ball grid array-module (BGA-module) incor- 67 porating separate Tx and Rx antennas (to avoid a lossy switch 68 at 60 GHz in TDD mode) integrated with a shaped 3-D-printed 69 plastic lens is proposed for integration with a smart eyewear 70 device for high-speed video transfer from the device to a lap- 71 top or a TV in front of the user. The transceiver is based on an 72 RFIC design using 65 nm CMOS technology and aims to ful- 73 fill the WiGig requirements for its highest available data rate, 74 i.e., MCS20. This mode offers a 4.158-Gbps data rate thanks 75 to OFDM 16-QAM modulation. All WiGig frequency sub- 76 bands are covered from 57 to 66 GHz. The chipset is described 77 in detail in [7]. The shaped lens is intended to achieve an 78 acceptable gain(> 10 dBi) and to shape the radiation pattern 79 0018-926X © 2016 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission.

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with wide beamwidth in the horizontal plane (at least 100◦ at

80

5 dBi gain) and narrow beamwidth in the vertical plane (in the 81

order of 25◦ at 5 dBi gain). The challenge for its design is 82

the additional need to counteract the beam depointing effect 83

due to the impossibility of positioning the separate Tx and 84

Rx radiating elements simultaneously at the lens single focal 85

point. The in-plane separation of these radiating elements is 86

6.9 mm. The lens was found to be a convenient solution to 87

address the shaped beam challenge instead of using a large 88

planar array with low aperture efficiency and nonuniform (and 89

lossy) feeding network. The objective of the paper is to show 90

that the proposed antenna concept is feasible for mm-wave 91

eyewear applications, being compact, low cost and with neg-92

ligible impact in terms of head specific absorption rate (SAR). 93

Section II gives some basic information about the BGA-module 94

and the Tx and Rx radiating elements. The design of the lens 95

with its theoretical background and simulation results are also 96

explained in this section. The integration of the BGA-module 97

and the lens within the eyewear is presented in Section III. 98

Simulation results taking into account the presence of the 99

user’s head are also presented in this section. Measurement 100

results for the manufactured prototype are given in Section IV. 101

Section V discusses the evaluation of the radiation exposure on 102

the body through simulations. Finally, conclusion is drawn in 103

Section VI. 104

II. ANTENNADESIGN 105

A. BGA-Module

106

The BGA-module was designed and manufactured in high 107

density integration (HDI) technology dedicated to 60 GHz SiP 108

solutions. This HDI technology is based on standard BGA 109

design and realization techniques: it enables a minimum trace 110

resolution as well as trace spacing of 50 μm. The low-cost 111

stack-up of three organic substrates enables four metallization 112

layers. A picture of the BGA-module can be seen in Fig. 1. 113

This module was designed to radiate in free-space. The mod-114

ule has equal length and width (12 × 12 mm2_{) with a height of} 115

0.5 mm. It hosts two printed antennas, one for receiving and one 116

for transmitting, offset from the center of the BGA-module and 117

separated byΔd = 6.9 mm distance from each other (1.38 λ0). 118

The antennas are of aperture-coupled patch type, where the 119

apertures are excited through a microstrip line underneath them. 120

The antennas are linearly polarized with a measured gain higher 121

than 4 dBi between 57 and 66 GHz (including transmission line 122

losses). More information about the BGA-module (version 1) 123

and the antenna can be found in [8]. They are thus not repeated 124

here for the sake of brevity. However, it should be noted that 125

a second optimized version of the BGA-module is used in the 126

current paper having more than 10 dB return loss and 5 dBi 127

gain from 57 to 66 GHz which is better than the performance 128

presented in [8]. 129

B. Lens Design

130

A small shaped lens is used directly on top of the patch ele-131

ments of BGA-module to modify its radiation pattern into a 132

Fig. 1. Picture of the BGA-module version 2: bottom view on left and top view on right.

F1:1 F1:2

Fig. 2. Schematic view of the optimized 3-D-lens placed above the BGA-module with axis definition.

F2:1 F2:2

more convenient shape for the intended application. The gain 133 is required to exceed 5 dBi within a 100◦ angular interval in 134 the horizontal plane (H-plane, or yz-plane in Fig. 2) and 25◦ 135 angular interval in the vertical plane (E-plane or xz-plane in 136 Fig. 2). This gives the user enough margin to look comfort- 137 ably at the screen from different angles without compromising 138 the data link. It should be noted that no codesign between the 139 source and the lens was achieved as we reused the existing 140 BGA-module version 2 dedicated to radiate in free space (and 141

not plastic medium). 142

The radiation pattern described in the paragraph above is 143 not symmetric and therefore requires a full 3-D shaped lens. 144 Although geometrical optics (GO) formulations exist for the 145 design of arbitrary shaped dielectric lenses subject to arbitrarily 146 output power template conditions [9]–[10], it is shown in this 147 study that for the present radiation pattern specifications it is 148 enough to consider a far simpler and computationally fast alter- 149 native based on a modification of the GO formulation for axial 150

symmetric lenses. 151

The combination with physical optics (PO) analysis enables 152 faster optimization of the shape of the 3-D lens compared 153 to the exact 3-D GO formulation, offering a very reasonable 154 agreement with the targeted radiation pattern beamwidths. 155 The GO/PO-based lens design procedure requires prior 156 knowledge of the radiation pattern of one antenna of the BGA- 157 module into an unbound medium of the chosen material for 158

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Fig. 3. Normalized simulated E- and H-plane radiation patterns of the BGA-module (Tx-antenna) in an unbounded medium of ABS plastic. Symmetric RP is an average of the E- and H-planes further used for the lens design. F3:1

F3:2 F3:3

the lens. ABS-M30 plastic material (consumer grade plastic 159

used for smartphone casing) was chosen in order to ensure 160

low cost for the overall system. A 3-D-printing rapid man-161

ufacturing technology was selected to fabricate the lens. A 162

disk sample of ABS material was printed to experimentally 163

evaluate its complex permittivity. The Fabry–Perot resonator 164

measurement method presented in [11] gave usr= 2.48 and

165

tan(δ) = 0.009 at 60 GHz. The lens design was performed at 166

60 GHz, as the central frequency. Intrinsic to the GO design, 167

the frequency bandwidth of the lens is inherently large but 168

the full-system bandwidth is mainly determined by the BGA-169

module bandwidth. The radiation pattern of the feed inside 170

the unbounded ABS medium at 60 GHz was obtained from a 171

full-wave HFSS simulation (Fig. 3). Overall, the main E- and 172

H-planes of the bare BGA antennas have similar beamwidths. 173

The 3-D shaped lens is designed in two steps. First, the lens 174

profile in the horizontal plane is obtained from an elevation cut 175

of an axial-symmetric lens designed with an appropriate GO 176

formulation. Then, the design rule for the lens profile in the ver-177

tical plane is defined and the complete 3-D lens physical shape 178

is obtained from an adequate combination of both horizontal 179

and vertical lens profiles. 180

Fig. 4 shows the general geometry for the lens design pro-181

cedure. The axial symmetric lens profile is represented byr(η)

182

where r(η = 0◦) = 15 mm corresponds to the total height of

183

the lens. The feed, corresponding to one patch of the BGA-184

module, is assumed initially to be at the center of the base of the 185

lens and in direct contact with it. The lens design assumes that 186

the feed is positioned at the center of the lens. However, the tar-187

get radiation patternG(θ) that is needed to define the lens shape

188

is carefully optimized to minimize its dependence with feed off-189

set when the BGA-module is integrated at the base of the lens. 190

At this point, the lens design formulation requires that the feed 191

radiation patternU (η) is axial symmetric. The symmetric U (η)

192

power pattern is generated as an average of the cocomponents 193

in the main planes of the BGA antenna (symmetric RP curve in 194

Fig. 4. Horizontal (or H-plane) profile of the plastic lens. F4:1

Fig. 3). The obtainedU (η) function is represented in Fig. 3 by 195 the black dashed curve (symmetric RP). Other symmetrization 196 options could have been adopted for the feed radiation pattern. 197 They would imply differentr(η = 0) values and different G(θ) 198 shape from what we obtained in the optimization process. The 199 selected target patternG(θ) is a flat-top type with a sharp drop- 200 off atθ = 60◦to comply with the desired full 100◦beamwidth 201

in the horizontal plane of the lens. 202

The lens profile r(η) is designed to transform U (η) into 203 a target axial symmetric power pattern G(θ). The r(η) pro- 204 file is obtained by solving a set of two differential equations 205 defined by the authors in [12]. The integration is performed 206 for increasingη angles up to the ηmaxvalue where ∂r(η)/∂η 207 becomes negative. Thisηmaxangle defines the edge of the lens 208 (Fig. 4). The remaining points in the lens profile from z = 209

r(ηmax) cos(ηmax) to z = 0 are defined with a constant value 210 ofx = r(ηmax) sin(ηmax). By forcing nonnegative ∂r(η)/∂η, 211 the lens surface will not diffract the radiation of the feed into the 212 negativez direction, i.e., into the user. The analytical formula- 213 tion is valid for arbitrary initial valuesr(η = 0◦) which acts 214 as a scaling factor. In the axial symmetric lens, the matching 215 between the output power pattern and the horizontal plane tar- 216 get improves asr(η = 0◦) increases. This value influences the 217 vertical plane power pattern in a different way, as will be dis- 218 cussed ahead. In the present design,r(η = 0◦) = 15 mm was 219 chosen as a compromise between the output power pattern spec- 220 ification in both planes and the utmost size constraint for the 221 desired integration with the glasses. The obtained lens profile is 222 presented in Fig. 4. This curve is used as the horizontal profile 223 of the 3-D lens. In the vertical plane (E-plane), narrowing the 224 radiation pattern of the BGA-module can be achieved by using 225 a beam collimating lens profile (like an ellipse). For each cut 226 of the 3-D lens at a constantx value, an elliptical lens profile is 227 implemented (Fig. 5). Eachx-cut corresponds to a given η angle 228 so thatx = r(η) sin(η). In each x-cut, the height of the ellipti- 229 cal profile isF (η) = r(η) cos(η). The elliptical lens profile is 230

defined by 231 l (η, α) = √ εr− 1 √ εr− cos (α)F (η) (1)

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Fig. 5. Vertical cut of the plastic lens profile for a x= constant plane. F5:1

where α is the angle of each point l(η, α) in relation to the

232

vertical axis of each cut plane of the lens profile as indicated in 233

Fig. 5. Therefore, the complete 3-D lens profile is defined by 234

the following set of parametric equations 235

x (η, α) = r (η) sin (η) y (η, α) = l (η, α) sin (α)

z (η, α) = l (η, α) cos (α) . (2)

As with η, the α angle also ranges from 0 to αmax where 236

∂l(ηα)/∂α becomes negative. The remaining points from

237

z = z(η, αmax) to z = 0 are defined with a constant y = 238

y(η, αmax). 239

The obtained 3-D lens profile is shown in Fig. 2. Its over-240

all size isΔz = 15 mm by Δx = 26 mm and Δy = 14 mm. 241

This approximate 3-D lens design procedure is an evolution of 242

the one developed by the authors in [13]. The corresponding 243

radiation pattern is calculated using PO, considering the actual 244

nonsymmetric feed radiation pattern shown in Fig. 3. 245

The normalized result is presented in Fig. 6(a) for the main 246

planes, confirming the effectiveness of the proposed design. 247

The simulated maximum directivity is of the order of 12 dBi. 248

To achieve a narrower E-plane radiation pattern and a higher 249

lens directivity, the lens size can be increased by choosing a 250

higher value forr(η = 0◦).

251

Due to the Rx and TxΔd = 6.9 mm separation in the BGA-252

module, the lens is not fed from its focal point at the center of 253

the base of the lens. TheΔd/2 = 3.45 mm (0.69 λ₀at 60 GHz) 254

feed off-set in the x-axis tends to produce a beam

depoint-255

ing effect. However, the previous H-plane radiation pattern 256

template was specifically chosen to minimize this effect. It is 257

noted that they-plane elliptical profile does not allow

depoint-258

ing minimization if y-axis feed off-set was selected instead.

259

Thex-axis feed off-set effect in the horizontal plane (H-plane)

260

radiation pattern of the lens can be seen in Fig. 6(b), show-261

ing that the flat-top characteristic is reasonably maintained and 262

only 1 dB reduction is observed in the broadside direction from 263

the nonoffset source case. It has little influence in the E-plane 264

since Tx and Rx patches remain in the focal point of the ellip-265

tical x-cut profile of the lens that passes through each feed

266

position. 267

Fig. 6. Normalized GO/PO simulated radiation pattern of the 3-D lens fed by the Tx patch of the BGA-module at 60 GHz. (a) Feed at the center of the lens. (b) Feed offset from the center by x= 3.45 mm.

F6:1 F6:2 F6:3

III. INTEGRATION OF THEBGA-MODULEWITH THE 268

EYEWEARDEVICE 269

In order to validate the GO/PO-based lens design and to eval- 270 uate the antenna in the realistic use-case scenario, full wave 271 electromagnetic simulations were also carried out, using the 272 commercial software Empire XCcel [14]. The simulation model 273 included the BGA-module integrated with the dielectric lens, 274 mounted in the left-hand side of a dedicated ABS eyewear 275 frame (Fig. 7 left part). An FR4 substrate which might be 276 needed in a realistic product as the application PCB and a back- 277 ing ground plane was also included in this model, behind the 278 antenna-module (Fig. 8). The frame includes a curved region on 279 the right-hand side of the head to emulate visually the screen of 280 a smart eyewear device. It also includes on the two sides of the 281 frame, two parallelepiped casing-like structures for housing the 282 application PCBs for WLAN/Bluetooth and WiGig standards, 283

respectively. 284

A homogeneous specific anthropomorphic mannequin 285 (SAM) head was also included in the simulation to account for 286 the user head influence. Considering the computation time and 287 memory requirements to simulate the full set-up from Fig. 7, 288 it was decided to use a cropped model of the head, since the 289 effects of the tissues that are placed far from the antenna in 290 terms of wavelength will be negligible. The cropped model 291 used in the simulations can be seen in Fig. 7 (right side). It 292 keeps all the structures and materials that lie within10 λ₀dis- 293 tance from the Antenna-module and discards all the others, 294 so the final dimensions of the simulation rectangular box is 295 78 × 73 × 155 mm3_{. The lens and the dielectric frame were} ₂₉₆ modeled as ABS plastic material. The values taken from Fabry– 297 Perot measurements and given in Section II B were used to 298

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Fig. 7. Simulation model of the BGA-module integrated with the lens and the ABS eyewear frame close to the SAM Head. The shaped lens is on the frame left side.

F7:1 F7:2 F7:3

Fig. 8. Position of the FR4 substrate and the ground plane which is backing the antenna-module.

F8:1 F8:2

model the ABS material. The outer shell of the head (marked 299

as Tissue-1) was assigned the properties of dry skin at 60 GHz, 300

having a relative permittivity of 7.98 and a loss tangent of 1.37. 301

To model the interior region of the head (Tissue-2), electrical 302

properties of the brain was used, with a relative permittiv-303

ity of 10.4 and loss tangent of 1.19. Those values were taken 304

from [15]. 305

A second set of simulations was also performed removing 306

the head and the backing PCB of the lens. The comparison 307

of the simulated reflection coefficient of the Tx antenna of the 308

BGA-module with the lens without the head and PCB and with 309

the head and PCB is presented in Fig. 9. The simulated cou-310

pling coefficient from the Tx antenna to Rx antenna is very low 311

(|S21| < −30 dB) and thus is not shown here. The Tx antenna

312

integrated with the lens with the backing PCB and head has a 313

reflection coefficient always lower than −6.7 dB between 57 314

and 66 GHz, even decreasing below−15 dB around the lower 315

edge of the band. It can also be seen from the same figure that 316

the reflection coefficient of the same Tx antenna without back-317

ing PCB and head is very similar, suggesting negligible effects 318

of the head and the backing PCB. The consequence of the 319

absence of a codesign between the source and the lens directly 320

translates into a frequency shift (compared to the “without head 321

and PCB case”) with a minimum of|S₁₁| around 55–57 GHz 322

(almost out-of-band) as the BGA-module now radiates into 323

plastic rather than air. 324

Fig. 9. Simulated reflection coefficient for the Tx patch of the BGA-module integrated with the lens without the Head and PCB and with the Head and PCB.

F9:1 F9:2

Fig. 10. Full-wave simulated directivity patterns of the BGA-module integrated with the lens with head and PCB and without head and PCB. (a) E-plane and (b) H-plane at 60 GHz.

F10:1 F10:2 F10:3

The comparison of the full-wave radiation patterns in the 325 E-plane (ϕ = 0◦) and H-plane (ϕ = 90◦) for the two configura- 326 tions is shown in Fig. 10(a) and (b), respectively. The maximum 327 radiation does not occur exactly in the front direction of the eye- 328 wear. There is a slight asymmetry in the radiation pattern (10◦ 329 tilt) but this is not really important in this application as the 330 beam tilt is small as compared to the beamwidth and the user 331 does not necessarily have to be directly in front of the receiv- 332 ing device. Note that the obtained radiation patterns confirm 333 that the lens geometry is fairly suitable to overcome the focal 334 depointing. In addition, the specified 5 dBi gain beamwidth of 335 25◦in the E-plane and 100◦in the H-plane is very closely met. 336 The maximum full-wave simulated directivity is almost 15 dBi 337 which is higher than the 12 dBi simulated with GO/PO. We 338 anticipate that the difference is mainly the result of known lim- 339 itations of the GO/PO asymptotic method for small-size lenses. 340 Also, it is known that a surface wave may appear for ellipti- 341 cal lenses at the lens/air interface which, for some lens sizes, 342 can lead to higher directivity than predicted with GO/PO [16]. 343 The head and the PCB behind the antenna-module have little 344

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Fig. 11. Pictures of the manufactured ABS frame incorporating the shape of the lens.

F11:1 F11:2

Fig. 12. BGA-module and 3-D-printed ABS plastic lens. F12:1

effects (but not significant) in the general radiation pattern (1 dB 345

reduction of the directivity and broader beam is the E-plane); 346

however, we decided to omit this last configuration in the next 347

simulation studies and measurements. 348

IV. MEASUREMENTS 349

A. Fabrication of the Eyewear Prototype and the

Antenna-350

Module

351

A picture of the manufactured ABS frame integrating the lens 352

is shown in Fig. 11. Fig. 12 presents the BGA-module (left side) 353

and the ABS lens alone (right side). All the ABS prototypes 354

were fabricated using 3-D printing plastic technology. 355

B. Measurement Results

356

The measurements of the BGA-module with the integrated 357

lens were carried out without the head and PCB, following the 358

Fig. 13. Pictures of the probing of the Tx antenna of the BGA-module with the lens (left side), bottom view of the BGA-module with the lens inside the foam support (right side).

F13:1 F13:2 F13:3

Fig. 14. Simulated and measured reflection coefficient of the Tx antenna of the BGA-module with the lens.

F14:1 F14:2

Fig. 15. Simulated and measured 3-D realized gain patterns of the BGA-module with the lens at 60 GHz.

F15:1 F15:2

conclusion from the previous section. Also, the full eyewear 359 frame was not utilized for the radiation pattern measurements 360 as it physically impairs the access of the feeding probe used in 361 our millimeter-wave measurement set-up [17] (Fig. 13). 362 There is a good agreement between measured and simulated 363 reflection coefficients given in Fig. 14. The measured reflection 364 coefficient is well below −11.5 dB in the target band (57– 365 66 GHz). Note, no codesign was performed which suggests that 366 better performance could be achieved in a possible new version 367 of the eyewear and BGA-module. The simulated and measured 368 realized gain patterns are presented in 3-D form in Fig. 15 369 and in the main planes in Fig. 16(a) (E-plane forϕ = 0◦) and 370 Fig. 16(b) (H-plane forϕ = 90◦). The maximum measured gain 371 is approximately 11 dBi at 60 GHz (including the transmis- 372 sion line losses). The measured 5 dBi gain beamwidth is 24◦ 373 in E-plane and 96◦in H-plane. A comparison of the simulated 374 and measured radiation efficiency can be seen in Fig. 17. A fair 375 agreement is observed, especially in the target band between 376 57 and 66 GHz. The measured efficiency has been extracted 377 from the 3-D realized gain pattern with the method already 378 presented in [18]. The Tx antenna of the BGA-module has a 379 measured radiation efficiency ranging from 52% to 58% in this 380 band which is a suitable value for WiGig transmissions between 381

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Fig. 16. Simulated and measured realized gain of the BGA-module with the lens in (a) E-plane and (b) H-plane at 60 GHz.

F16:1 F16:2

Fig. 17. Simulated and measured radiation efficiency of the Tx antenna from the BGA-module with the lens.

F17:1 F17:2

the eyewear and a TV or a laptop. The stability of the gain pat-382

tern versus frequency was also investigated through the band 383

of interest. Fig. 18 presents the measured realized gain pat-384

terns (E- and H-planes) for three frequency points: 58, 60, and 385

66 GHz. These patterns show negligible variation with respect 386

to frequency even in terms of 5 dBi beamwidth as presented in 387

Table I, all complying with the beamwidth specification. 388

V. HUMANBODYEXPOSURE 389

International standards based on the incident power den-390

sity have been developed to limit the electromagnetic exposure 391

by the human body from RF devices at 60 GHz. The IEEE 392

(USA) recommends a maximum power density of 10 W/m2 393

averaged over0.01 m2(10 cm × 10 cm) averaged over 3.6 min 394

for the general public [19]. The limit is100 W/m2in controlled 395

environments averaged over 21.6 s [19]. The standards also 396

determine a maximum power density of1000 W/m2averaged 397

over any one square centimeter. ICNIRP (Europe) has power 398

density limits of, respectively,10 W/m2and50 W/m2averaged 399

over20 cm2 for the general public and controlled conditions. 400

Fig. 18. Measured realized gain in (a) E-plane and (b) H-plane of the Tx antenna of BGA-module with the lens for different frequencies.

F18:1 F18:2

TABLE I T1:1

COMPARISON OFMEASUREDBEAMWIDTH ATDIFFERENTFREQUENCIES T1:2

The maximum power density averaged over1 cm2 should not 401 exceed 20 times the above values. The averaging time can 402 be calculated by68/f1.05 = 0.92 min. The ICNIRP levels are 403 stricter for both the larger averaging areas and also the1 cm2 404 area. Therefore, compliance with ICNIRP guarantees compli- 405 ance with the IEEE recommendation. Despite the power density 406 being defined in the standards, there is no consistent evaluation 407 metric in the recent published papers. The local specific absorp- 408 tion rate (SAR) is examined in [20] and [21]. The 1-g SAR and 409 the power absorbed were discussed in [22]. The maximum elec- 410 tric field, power density, and local SAR were assessed in [23]. 411

In vitro protein and culture were considered in [24] and [25] 412 where the maximum local SAR and the SAR averaged over 413 the whole sample was related to the incident power density. A 414 thermal imaging camera was used to measure the temperature 415 distribution and hence the local and average power density as 416 well as the local SAR in [26]. This paper concluded that power 417 levels up to 550 mW would comply with the exposure limit 418 and an incident power density of 10 W/m2 would result in a 419

temperature increase of0.1◦C. 420

In our study, two sets of simulations were performed for ana- 421 lyzing the effect of the BGA-module with lens on the head 422 of the user. The first set of simulations was performed with a 423 homogeneous head model as used in Section III, consisting of 424 the outer part modeled as dry skin and the inner part modeled as 425 the brain tissue. The second set of simulations was performed 426

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IEEE Proof

Fig. 19. Simulated power density with SAM head (scaled to 1 W input power) over the skin for (a) 60 GHz and (b) 66 GHz.

F19:1 F19:2

with the visible human head model taking into account different 427

tissues with corresponding electrical properties. 428

A. Simulations With Homogeneous Head Model

429

Simulations with the homogeneous head model were per-430

formed to obtain the power density level at the surface of 431

the skin. According to ICNIRP guidelines, the power density 432

level on the tissue, averaged over20 cm2should be lower than 433

1 mW/cm2_(or_{10 W/m}2_{) around 60 GHz frequency. The} simu-434

lated power density with EMPIRE XCcel software on the skin 435

surface can be observed in Fig. 19 for 60 and 66 GHz. It should 436

be noted here that the values presented in this figure legends are 437

not averaged either over time or space, so they are the worst-438

case levels. Moreover, the input power to the antenna was set 439

as 1 W in the simulations so the power density levels shown 440

in this figure need to be divided by 100 to comply with the 441

typical input power level of 10 dBm for WiGig devices. The 442

unaveraged maximum power density observed on the skin sur-443

face varies between 0.37 and0.42 W/m2between 60–66 GHz 444

(for an input power of 10 dBm). These values are well below 445

the reference values from the guidelines even though they are 446

Fig. 20. Simulated power density with VH head (scaled to 1 W input power) for (a) 60 GHz and (b) 66 GHz.

F20:1 F20:2

instantaneous, unaveraged values. Obtaining such a low-power 447 density over the skin was expected since the main radiation 448 is directed away from the head and also the spacing between 449 the antenna-module and the head is 30 mm which is6 λ0 at 450

60 GHz. 451

B. Simulations With Visible Human Head Model 452

The simulations were also performed by placing the BGA- 453 module with lens on a truncated VH model with 1 mm res- 454 olution (inset in Fig. 20). The Yee cell size inside the head 455 was set to 0.2 mm and even smaller cells were used to dis- 456 cretize the BGA-module including the lens. Empire can display 457 the power density on the surface of the SAM head, which is 458 categorized as a solid shape, but not on the voxel-based VH 459 head. Therefore, to examine the power density on the surface 460 of the VH head, the SAM-shaped field monitor was copied to 461

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IEEE Proof

head. Therefore, the SAM-shaped field monitor was moved

sev-464

eral millimeters toward the antenna to lie on the surface of the 465

VH head. Approximate calculations using the path loss equa-466

tion indicated that this positioning difference can increase the 467

power density by approximately 5 dB. The power density on 468

the surface of the VH head varies between 67 and 100 W/m2 469

as shown in Fig. 20. These values are again the maximum val-470

ues seen at a point in space, normalized to 1 W input power, 471

without any space or time averaging. Considering 10nbsp;dBm 472

of maximum input power in WiGig applications, the values are 473

well below the standard values, even without averaging. 474

VI. CONCLUSION 475

This paper demonstrated the feasibility of a compact low-476

cost antenna assembly for a WiGig smart eyewear device 477

intended for high-speed wireless data communication in the 478

60-GHz band with a laptop or TV facing the user. The antenna-479

module incorporates a 3-D-printed shaped dielectric lens espe-480

cially designed to enhance gain while shaping the radiation 481

pattern to provide wide angular coverage in the horizontal plane 482

and narrow beam coverage in the vertical plane. The lens design 483

was based on GO/PO but full wave electromagnetic simulation 484

was carried out to evaluate the antenna assembly performance 485

when integrated with the eyewear device. Results were pre-486

sented for two scenarios: the first one included the user’s head 487

as well as a portion of a PCB backing the antenna-module. 488

In the second scenario, the head and PCB were removed. It 489

was demonstrated that the effect of the head and of the back-490

ing PCB on the radiation pattern and the reflection coefficient 491

was negligible. Keeping this in mind, the measurements for the 492

BGA antenna-module and the lens were carried out in free-493

space, showing a good agreement with the simulations in terms 494

of reflection coefficient, radiation efficiency, and realized gain 495

radiation pattern. The maximum measured gain at 60 GHz was 496

11 dBi, with 5 dBi gain beamwidth of 24◦in the vertical plane 497

and 96◦in the horizontal plane (including the transmission line 498

losses). The measured gain radiation pattern was also shown 499

to have negligible variation over the target frequency band 500

(57–66 GHz). 501

The effects of the antenna radiation on the human body was 502

analyzed for two sets of simulations, using both a homogeneous 503

SAM head phantom and visible human head model. The sim-504

ulated power density values for both head models were found 505

to be lower than the limits established in the related standards, 506

considering an input power of 10 dBm. 507

ACKNOWLEDGMENT 508

The authors would like to acknowledge Orange Labs La 509

Turbie and the CIM-PACA design platform. F. Devillers from 510

Orange Labs is deeply acknowledged for the adjustments of the 511

mechanical parts of the measurement set-up. They would also 512

like to acknowledge a scholarship given by project ESF RNP 513

“Newfocus” to Aimeric Bisognin. 514

[1] Google Glass [Online]. Available: http://www.google.com/glass/start/ 516Q2 [2] M100 Smart Glass [Online]. Available: http://www.vuzix.com/UKSITE/ 517

consumer/products_m100.html 518

[3] RECON Jet [Online]. Available: http://www.reconinstruments.com/ 519

products/jet/ 520

[4] Olympus MEG4.0 [Online]. Available: http://www.olympus.co.jp/jp/ 521 news/2012b/nr120705meg40j.jsp 522 [5] Optinvent ORA [Online]. Available: http://optinvent.com/see-through- 523

glasses-ORA 524

[6] A. Cihangir et al., “Feasibility study of 4G cellular antennas for eyewear 525 communicating devices,” IEEE Antennas Wireless Propag. Lett., vol. 12, 526

pp. 1704–1707, 2013. 527

[7] A. Siligaris et al., “A 65 nm CMOS fully integrated transceiver mod- 528 ule for 60 GHz wireless HD applications,” IEEE J. Solid-State Circuits, 529 vol. 46, no. 12, pp. 162–164, Dec. 2011. 530 [8] R. Pilard et al., “HDI organic technology integrating built-in antennas 531 dedicated to 60 GHz SiP solution,” in Proc. IEEE Antennas Propag. Soc. 532

Int. Symp. (AP-SURSI), Chicago, IL, USA, Jul. 2012, pp. 8–14. 533 [9] C. Salema, C. A. Fernandes, and R. K. Jha, Solid Dielectric Horn 534

Antennas. Norwood, MA, USA: Artech House, 1998. 535 [10] R. Sauleau and B. Bares, “A complete procedure for the design and 536 optimization of arbitrarily shaped integrated lens antennas,” IEEE Trans. 537

Antennas Propag., vol. 54, no. 4, pp. 1122–1133, Apr. 2006. 538 [11] C. A. Fernandes and J. R. Costa, “Permittivity measurement and 539 anisotropy evaluation of dielectric materials at millimeter-waves,” in 540

Proc. IMEKO World Congr., Lisboa, Portugal, Sep. 2009. 541Q3 [12] C. A. Fernandes, “Shaped dielectric lenses for wireless millimeter-wave 542

communications,” IEEE Antennas Propag. Mag., vol. 41, no. 5, pp. 141– 543

150, Oct. 1999. 544

[13] C. A. Fernandes, “Design of shaped lenses for non-symmetric cells in 545 MBS,” in Proc. IEEE Antennas Propag. Soc. Int. Symp. (AP-S/URSI), 546 Orlando, FL, USA, Jul. 1999, vol. 3, pp. 2440–2443. 547 [14] EMPIRE XCcel [Online]. Available: http://www.empire.de/ 548 [15] [Online]. Available: http://www.itis.ethz.ch/itis-for-health/tissue- 549

properties/database/tissue-frequency-chart/ 550Q4

[16] D. Pasqualini and S. Maci, “High-frequency analysis of integrated dielec- 551 tric lens antennas,” IEEE Trans. Antennas Propag., vol. 52, no. 3, 552

pp. 840–847, Mar. 2004. 553

[17] D. Titz, F. Ferrero, and C. Luxey, “Development of a millimeter-wave 554 measurement setup and dedicated techniques to characterize the match- 555 ing and radiation performance of Probe-fed Antennas,” IEEE Antennas 556

Propag. Mag., vol. 54, no. 4, pp. 188–203, Aug. 2012. 557 [18] D. Titz, F. Ferrero, P. Brachat, C. Luxey, and G. Jacquemod, “Efficiency 558 measurement of probe-fed antennas operating at millimeter-wave fre- 559 quencies,” IEEE Antennas Wireless Propag. Lett., vol. 11, pp. 1194–1197, 560

Oct. 2012. 561

[19] IEEE Standard for Safety Levels With Respect to Human Exposure to 562

Radio Frequency Fields 3 kHz to 300 GHz, IEEE Standard C95.1-2005, 563

pp. 1–250, 2006. 564

[20] N. Chahat, C. Leduc, M. Zhadobov, and R. Sauleau, “Antennas and 565 Interaction with the Body for Body Centric Wireless Communications 566 at Millimeter Waves,” in Proc. Eur. Conf. Antennas Propag. (EuCAP), 567

2013, pp. 772–775. 568

[21] K. H. Chan, S. W. Leung, Y. L. Diao, Y. M. Siu, and K. T. Ng, “Analysis 569 of millimeter wave radiation to human body using inhomogeneous mul- 570 tilayer skin model,” in Proc. Asia-Pac. Symp. Electromagn. Compat. 571

(APEMC), 2012, pp. 721–724. 572

[22] A. R. Guraliuc, M. Zhadobov, G. Valerio, and R. Sauleau, “Enhancement 573 of on-body propagation at 60 GHz using electro textiles,” IEEE Antennas 574

Wireless Propag. Lett., vol. 13, pp. 603–606, Apr. 2014. 575 [23] M. Kouzai, A. Nishikata, T. Sakai, and S. Watanabe, “Characterization 576 of 60 GHz millimeter-wave focusing beam for living-body exposure 577 experiments,” in Proc. EMC’09, 2009, pp. 309–312. 578 [24] M. Zhadobov et al., “Evaluation of the potential biological effects of 579 the 60-GHz millimeter waves upon human cells,” IEEE Trans. Antennas 580

Propag., vol. 57, no. 10, pp. 2949–2956, Oct. 2009. 581 [25] M. Zhadobov, R. Sauleau, R. Augustine, C. Le Quément, Y. Le Dréan, 582 and D. Thouroude, “Near-field dosimetry for in vitro exposure of human 583 cells at 60 GHz,” Bioelectromagnetics, vol. 33, no. 1, pp. 55–64, Jan. 584

2012. 585

[26] N. Chahat, M. Zhadobov, L. Le Coq, S. I. Alekseev, and R. Sauleau, 586 “Characterization of the interactions between a 60-GHz antenna and the 587 human body in an off-body scenario,” IEEE Trans. Antennas Propag., 588 vol. 60, no. 12, pp. 5958–5965, Dec. 2012. 589

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IEEE Proof

590 Aimeric Bisognin (S’xx) was born in Toulouse,

Q5 590

France, in 1989. He received the Engineering 591

degree in electronics from Polytech’Nice Sophia, 592

Sophia Antipolis, France, in 2012, and the M.S. 593

degree in telecommunications from EDSTIC, Sophia 594

Antipolis, France, in 2012, and the Ph.D. degree 595

in electronics engineering (with Hons.) from the 596

University of Nice Sophia-Antipolis, France, in 2015. 597

During his Ph.D., he worked with the Electronique 598

pour Objets Communicants (EpOC) Laboratory, 599

STMicroelectronics, Crolles, France. He is currently 600

a Postdoctoral Researcher with the EpOC Laboratory. He has authored or 601

coauthored 8 publications in journals and 19 publications in international 602

conferences. His research interests include millimeter-wave communications, 603

especially in the field of the design and measurement of antenna in package, 604

lens and reflector antennas for the 60, 80, and 120 GHz frequency bands. 605

606 Aykut Cihangir was born in Ankara, Turkey, in

606

1985. He received the Bachelor’s degree in electrical 607

and electronics engineering and the Master’s degree

Q6 608

from Middle East Technical University (METU), 609

Ankara, Turkey, in 2010, and the Ph.D. degree from 610

the University of Nice-Sophia Antipolis (UNSA), 611

Nice, France, in 2007, 2010, and 2014, respectively. 612

From 2007 to 2011, he was with the Aerospace

Q7 613

Industry as a Communication Systems Design 614

Engineer in Unmanned Aerial Vehicle projects. Then, 615

he was as a Postdoctoral Researcher with the EpOC 616

team, UNSA. His research interests include electrically small antennas, match-617

ing networks, and mobile terminal antennas. 618

619 Cyril Luxey (SM’xx) was born in Nice, France, in

619

1971. He received the Master’s degree (DEA) and 620

the Ph.D. degree in electrical engineering (both with 621

Hons.) from the University Nice-Sophia Antipolis, 622

Nice, France, in 1996 and 1999, respectively. 623

During his thesis, he worked on printed leaky-624

wave antennas, quasi-optical mixers, and retrodirec-625

tive transponders. From 2000 to 2002, he was with 626

Alcatel, Mobile Phone Division, Colombes, France, 627

where he was involved in the design and integration 628

of internal antennas for commercial mobile phones. 629

In 2003, he was an Associate Professor with the Polytechnic School, University 630

Nice Sophia-Antipolis. Since 2009, he has been a Full Professor with the IUT 631

Réseaux et Télécoms in Sophia-Antipolis, Valbonne, France. He is doing his 632

research in the EpOC Laboratory as the Vice-Deputy of this laboratory. In 633

October 2010, he was a Junior Member of the Institut Universitaire de France 634

(IUF). Also, he collaborates with Berkeley Wireless Research Center, Berkeley, 635

CA, USA, and Stanford University, Stanford, CA, USA, on mm-wave front-end 636

transceivers at mm-wave frequencies. He also works on electrically small anten-637

nas, multiantenna systems for diversity, and MIMO techniques. He has authored 638

or coauthored more than 250 papers in refereed journals, in international and 639

national conferences and as book chapters. He has given more than 10 invited 640

talks. His research interests include the design and measurement of millimeter-641

wave antennas, antennas-in-package, plastic lenses, and organic modules for 642

60, 120, and 240 GHz frequency bands. 643

Dr. Luxey is an Associate Editor for the IEEE ANTENNAS ANDWIRELESS 644

PROPAGATIONLETTERS, a Regular Reviewer for several IEEE and IET jour-645

nals and several European and US conferences in the field of microwave, micro-646

electronics, and antennas. He has been the General Chair of the Loughborough 647

Antennas and Propagation Conference 2011, the Award and Grant Chair 648

of EuCAP 2012, and the Invited Paper CoChair of EuCAP 2013. He will 649

be the TPC Chair of the EuCAP 2017 conference in Paris. He is also the 650

Delegate of the French Research Ministry for the COST IC1102 action VISTA 651

(Versatile, Integrated, and Signal-aware Technologies for Antennas) within the 652

ICT Domain. He and his students received the H.W. Wheeler Award of the 653

IEEE Antennas and Propagation Society for the best application paper of the 654

year 2006. He was also the corecipient of the Jack Kilby Award 2013 of the 655

ISSCC conference and the Best Paper Award of the EUCAP2007 conference, 656

the Best Paper Award of the International Workshop on Antenna Technology 657

(iWAT2009), the Best Paper Award at LAPC 2012, the Best Student Paper 658

Award at LAPC 2013 (third place), the Best Paper Award of the ICEAA 2014 659

conference, and the Best Paper Award of the innovation contest of the iWEM 660

2014 conference (second place). 661

662

Gilles Jacquemod (M’xx) graduated from ICPI 662 (CPE) Lyon, and received the M.Sc. degree (DEA) in 663 microelectronics from Ecole Centrale Lyon, Écully, 664 France, in 1986, the Ph.D. degree in integrated elec- 665 tronics from INSA Lyon, Lyon, France, in 1989. 666 From 1990 to 2000, he was with LEOM, Ecole 667 Centrale Lyon, as an Associate Professor, on ana- 668 log integrated circuit design and behavioral model- 669 ing of mixed domain systems. In 2000, he joined 670 the LEAT Laboratory, Nice, France, the Ecole 671 Polytechnique, Palaiseau, France, and Nice-Sophia 672 Antipolis University, Nice, France, as a Full Professor. Since 2010, he has 673 been with EpOC Research Team (URE UNS). He is the author and coauthor of 674 more than 250 journal and conference papers, and holds 3 patents. His research 675 interests include analog integrated circuit design and behavioral modeling of 676 mixed domain systems. He is also involved in RF design applied to wireless 677

communication. 678

Dr. Jacquemod is the Director of the CNFM PACA pole. He was the President 679 of the CIM-PACA Design Platform. 680

681

Romain Pilard (M’xx) received the B.S. and 681 M.S. degrees in electronics engineering from 682 Polytech’Nantes, University of Nantes, Nantes, 683 France, in 2006, and the Ph.D. degree in electrical 684 engineering from Telecom Bretagne, Brest, France, 685

in 2009. 686

Since 2010, he has been with the 687 STMicroelectronics, Crolles, France, where he 688 works on the development of integrated antennas and 689 high-performance passive components in advanced 690 bulk and SOI RF CMOS technologies. His research 691 interests include millimeter-wave antenna design and packaging technology 692

development. 693

694

Frédéric Gianesello (M’xx) received the B.S. 694 and M.S. degrees in electronics engineering from 695 the Institut National Polytechnique de Grenoble, 696 Grenoble, France, in 2003 and the Ph.D. degree 697 in electrical engineering from the Joseph Fourier 698 University, Grenoble, France, in 2006. 699 He is currently working for STMicroelecetronics, 700 Crolles, France, where he leads the team responsi- 701 ble for the development of electromagnetic devices 702 (inductor, balun, transmission line, and antenna) 703 integrated on advanced RF CMOS/BiMOS (down 704 to 14 nm), silicon photonics, and advanced packaging technologies (3-D 705 Integration, FOWLP, etc.). He has authored and coauthored more than 110 706 refereed journal and conference technical articles. 707 Dr. Gianesello has served on the TPC for the International SOI Conference 708 from 2009 up to 2011 and he is currently serving on the TPC for the 709 Loughborough Antennas and Propagation Conference (LAPC). 710

711

Jorge R. Costa (S’97–M’03–SM’09) was born in 711 Lisbon, Portugal, in 1974. He received the Licenciado 712 and Ph.D. degrees in electrical and computer engi- 713 neering from the Instituto Superior Técnico (IST), 714 Technical University of Lisbon, Lisbon, Portugal, in 715 1997 and 2002, respectively. 716 He is currently a Researcher with the Instituto 717 de Telecomunicações, Lisbon, Portugal. He is also 718 an Associate Professor with the Departamento de 719 Ciências e Tecnologias da Informação, Instituto 720 Universitário de Lisboa (ISCTE-IUL), Lisbon, 721 Portugal. He is the coauthor of four patent applications and more than 100 722 contributions to peer reviewed journals and international conference proceed- 723 ings. More than twenty of these papers have appeared in IEEE Journals. His 724 research interests include lenses, reconfigurable antennas, MEMS switches, 725

UWB, MIMO, and RFID antennas. 726

Prof. Costa is currently serving as an Associate Editor for the IEEE 727 TRANSACTIONS ON ANTENNAS AND PROPAGATIONand he was a Guest 728 Editor of the Special Issue on “Antennas and Propagation at mm- and Sub mm- 729 Waves,” from the IEEE TRANSACTIONS ONANTENNAS ANDPROPAGATION, 730 April 2013. He was the Co-Chair of the Technical Program Committee of the 731 European Conference on Antennas and Propagation (EuCAP 2015) in Lisbon. 732

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IEEE Proof

cal and computer engineering from Instituto Superior 735

Técnico (IST), Technical University of Lisbon, 736

Lisbon, Portugal, in 1980, 1985, and 1990, respec-737

tively. 738

He joined IST in 1980, where he is currently a 739

Full Professor with the Department of Electrical and 740

Computer Engineering in the areas of microwaves, 741

radio wave propagation and antennas. He is a Senior 742

Researcher with the Instituto de Telecomunicações 743

and member of the Board of Directors. He has coauthored a book, a book chap-744

ter, more than 150 technical papers in peer-reviewed international journals and 745

conference proceedings and 7 patents in the areas of antennas and radiowave 746

propagation modeling. His research interests include dielectric antennas for 747

millimeter wave applications, antennas, and propagation modeling for personal 748

communication systems, RFID and UWB antennas, artificial dielectrics, and 749

metamaterials. 750

Dr. Fernandes was a Guest Editor of the Special Issue on “Antennas and 751

Propagation at mm- and Sub mm-Waves,” from the IEEE TRANSACTIONS ON 752

ANTENNAS ANDPROPAGATION, April 2013. 753

754 Eduardo B. Lima (S’14) received the Licenciado

754

and M.Sc. degrees in electrical and computer engi-755

neering from the Instituto Superior Técnico (IST), 756

Technical University of Lisbon, Lisbon, Portugal, in 757

2003 and 2008, respectively. He is currently working 758

toward the Ph.D. degree at IST. 759

He is also currently a Researcher with the Instituto 760

de Telecomunicações (IT), Lisbon, Portugal. Since 761

2004, he has been also responsible for development 762

and maintenance of all measurement control systems 763

at the RF laboratories. He is the coauthor of one 764

patent application and 32 technical papers in international journals and confer-765

ence proceedings in the area of antennas. One of which was awarded the CST 766

University Publication Award 2010. His research interests include dielectric 767

lens antennas, Fabry–Pérot cavity antennas, and transmit arrays. 768

Dr. Lima is currently serving as a Technical Reviewer for the IEEE 769

TRANSACTIONS ON ANTENNAS AND PROPAGATION (TAP) and he also 770

served as a TPC member in the EuCAP2015 conference. 771

tronic and electrical engineering and the Ph.D. 774 degree in electronic and electrical engineering from 775 Loughborough University, Loughborough, U.K., in 776 2000 and 2005, respectively. 777 Since 2007, he has been a Lecturer in Photonic 778 Systems with the Wolfson School of Mechanical, 779 Manufacturing, and Electrical Engineering, 780 Loughborough University. His research interests 781 include optically activated semiconductor switches 782 for microwave circuits, devices and antennas. He also conducts research on 783 specific absorption rates due to mobile communications devices and in-body 784

antennas. 785

Since 2011, he is the Coordination Chair of the Loughborough Antennas and 786

Propagation Conference series. 787

788

William G. Whittow (M’12–SM’12) received the 788 B.Sc. degree in physics and the Ph.D. degree in 789 computational electromagnetics from the University 790 of Sheffield, Sheffield, U.K., in 2000 and 2004, 791

respectively. 792

From 2004 to 2012, he was a Research Associate 793 with the School of Electronic, Electrical and 794 Systems Engineering, Loughborough University, 795 Loughborough, U.K. He was a Lecturer in Electronic 796 Materials Integration, University of Loughborough, 797 in 2012 and he became a Senior Lecturer in May 798 2014. He has authored more than 150 peer-reviewed journal and conference 799 papers in topics related to electromagnetic materials, synthetic dielectrics, 800 wearable antennas, VHF antennas, specific absorption rate, FDTD, bioelec- 801 tromagnetics, meta-materials, heterogeneous substrates, embroidered anten- 802 nas, inkjet printing, electromagnetic compatibility, phantoms and Genetic 803

Algorithms. 804

Dr. Whittow was the Coordinating Chair of the Loughborough Antennas 805 and Propagation Conference (LAPC). He is an Associate Editor of Electronics 806

Letters. He serves on the technical programme committees of several IEEE 807 international conferences. He has been asked to give 6 invited conference 808 presentations and a 4-day invited workshop on bioelectromagnetics. 809

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IEEE Proof

Q1: Please provide postal code for all the affiliations.

Q2: Please provide author name and year of publication for Refs. [1]–[5], and [14]. Q3: Please provide page range for Ref. [11].

Q4: Please provide complete details for Ref. [15].

Q5: Please provide membership history (year) for authors “Aimeric Bisognin, Cyril Luxey, Gilles Jacquemod, Romain Pilard, Frédéric Gianesello, and Chinthana J. Panagamuwa.”

Q6: Please provide field of study for Master’s and Ph.D. degrees of author “Aykut Cihangir.” Q7: Please provide location for Aerospace Industry in the biography section.

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IEEE Proof

Lens for WiGig Applications in Eyewear Devices

2

Aimeric Bisognin, Student Member, IEEE, Aykut Cihangir, Cyril Luxey, Senior Member, IEEE,

Gilles Jacquemod, Member, IEEE, Romain Pilard, Member, IEEE, Frédéric Gianesello, Member, IEEE,

Jorge R. Costa, Senior Member, IEEE, Carlos A. Fernandes, Senior Member, IEEE,

Eduardo B. Lima, Student Member, IEEE, Chinthana J. Panagamuwa, Member, IEEE,

and William G. Whittow, Senior Member, IEEE

3

4 5 6 7

Abstract—A ball grid array-module (BGA-module)

incorpo-8

rating a low-cost shaped dielectric lens is proposed for wireless

9

communications in the 60-GHz WiGig band between a smart

eye-10

wear, where it is integrated and facing a laptop or TV. The module,

11

which is codesigned with a 60-GHz transceiver, consists of two

12

separate identical antennas for transmitting (Tx) and receiving

13

(Rx). The in-plane separation of these elements is 6.9 mm both

14

being offset from the lens focus. This poses a challenge to the lens

15

design to ensure coincident beam pointing directions for Rx and

16

Tx. The shaped lens is further required to narrow the angular

17

coverage in the elevation plane and broaden it in the horizontal

18

plane. A 3-D-printed eyewear frame with an integrated lens and

19

a recess for proper BGA-module integration is fabricated in

ABS-20

plastic material. Measurements show a reflection coefficient below

21

−12 dB in the 57–66 GHz band. A maximum gain of 11 dBi is

22

obtained at 60 GHz, with 24◦ and 96◦ beamwidth at 5-dBi gain,

23

respectively, in the vertical and horizontal planes. The radiation

24

exposure is evaluated for a homogeneous SAM head phantom and

25

a heterogeneous visible human head. The simulated power

den-26

sity values for both models are found to be lower than the existing

27

standards.

28

Index Terms—60 GHz, antenna-in-package, eyewear, lens

29

antennas, plastic packaging, WiGig.

30

I. INTRODUCTION 31

W

ITH the never-ending improvement of the capabilities Q1 32

of wireless communication devices, the most critical

33

necessity has been to supply the user with higher and higher

34

data rates. This has led to both the improvement of the existing

35