Hindawi Publishing Corporation
International Journal of Antennas and Propagation
Volume 2014, Article ID 471935, 9 pages
http://dx.doi.org/10.1155/2014/471935

Research Article
Fabrication of a 77 GHz Rotman Lens on a High Resistivity
Silicon Wafer Using Lift-Off Process
Ali Attaran and Sazzadur Chowdhury
Department of Electrical and Computer Engineering, University of Windsor, CEI 401 Sunset Avenue, Windsor,
ON, Canada N9B 3P4
Correspondence should be addressed to Ali Attaran; ali1111@uwindsor.ca
Received 4 December 2013; Accepted 4 April 2014; Published 4 May 2014
Academic Editor: Young Joong Yoon
Copyright © 2014 A. Attaran and S. Chowdhury. This is an open access article distributed under the Creative Commons Attribution
License, which permits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited.
Fabrication of a high resistivity silicon based microstrip Rotman lens using a lift-off process has been presented. The lens features
3 beam ports, 5 array ports, 16 dummy ports, and beam steering angles of ±10 degrees. The lens was fabricated on a 200 𝜇m thickhigh resistivity silicon wafer and has a footprint area of 19.7 mm × 15.6 mm. The lens was tested as an integral part of a 77 GHz radar
where a tunable X band source along with an 8 times multiplier was used as the RF source and the resulting millimeter wave signal
centered at 77 GHz was radiated through a lens-antenna combination. A horn antenna with a downconverter harmonic mixer wasused to receive the radiated signal and display the received signal in an Advantest R3271A spectrum analyzer. The superimposedtransmit and receive signal in the spectrum analyzer showed the proper radar operation confirming the Rotman lens design.

1. Introduction
Frequency-Modulated Continuous-Wave (FMCW) radar
sensors based on analog or digital beamforming technology
have been proven to be effective for avoidance of automotive
collision or mitigation of collision damage. Investigationshows that instead of using microelectronic based analog ordigital beamforming engines, if Rotman lens type passive
beamformers can be used, the signal processing time as wellas power consumption, size, integration complexity, and system building cost will be reduced significantly [1–5]. Several
design approaches are available that propose using Rotman
lens for automotive radar applications. However, almost allof them depend on conventional fabrication technique [2]with expensive mass production. Recent investigation showsthat high resistivity silicon is an attractive choice to realize
high performance millimeter wave MMICs due to lower
attenuation and lower cost [4–6]. Additionally, the siliconprocessing technology is highly matured and the lenses canbe batch fabricated like conventional ICs to lower the costof automotive radars. In [4], a 60 GHz silicon based Rotman
lens has been presented. In this context, this paper presents
the design procedure of the Rotman lens at mm-frequency

range followed by 3D full wave simulation results. Proposed
fabricated and tested silicon based Rotman lens is a potential
to use in a low cost 77 GHz automotive radar.

2. Rotman Lens Design
Rotman lens is a passive beamforming device and the beam
can be steered in a specified direction by selective excitationof any one of the input ports (beam ports) of the lens as shownin Figure 1. Detailed design principles of Rotman lens are wellestablished [4, 7, 8] and can be summarized in 3 major steps.
Step 1. The lens geometrical specifications such as positionsof the beam ports on the focal arc and the array ports on
the inner contour are determined from the theory of ray
optics. Detailed procedure to determine the initial geometry
following the ray optics theory is available in [7].
Step 2. Once the initial geometry is determined, the twodimensional aperture theory of classical antenna analysis isapplied to determine electrical characteristics of the lens suchas the return loss, insertion loss, port isolation, and power
efficiency.

2International Journal of Antennas and Propagation
Radiating element
Array contour
Wave front

Beam port

Beam contour
Transmission line

Outer contour

Figure 1: A conceptual diagram of a Rotman lens.

Step 3. Finally, a 3D field solver like HFSS or XFDTD
is used to carry out a full-wave analysis that extends thetwo-dimensional analysis to three-dimensional analysis toincorporate the effects of reflections and scattering fromthe sidewalls (spillover power) and the lens cavity, surface
wave propagation, conductivity of the metal layer, conductive attenuation coefficient, and the dielectric attenuation
coefficient [6]. The initial specifications determined in Step 1
are more or less modified at this step to obtain optimum
performance.
Following this procedure, for the target automotive application, the lens center frequency has been selected as 77 GHz
for use in an automotive radar. For use in a trimode (shortrange, mid-range, and long-range) FMCW radar, the lens
bandwidth has been selected to be 8 GHz to match the
maximum chirp bandwidth of the FMCW signal [5].
Following [9], a scan angle of 10∘ has been selected tocover an azimuthal angle of 30∘ at −6 dB of antenna radiation
pattern. The number of array port has been chosen to be
5 to minimize the lens width. Following [7], the distance
between array elements is calculated to be 0.65𝜆. This element
spacing plays an important role in determining beam widthof the microstrip patch antenna arrays. This gives a length of
5.2 mm for 𝑁max which is the farthest output port in the outer
contour. Instead of normalizing the parameters with respect
to the focal length, following the method presented in [8], theparameters are normalized with respect to the total length of
the array aperture:𝐹 = 2𝑓𝑁max

sin (𝛽),sin (𝛼)(1)where 𝛽 is the angular direction of the wave front associatedwith the perfect focal point at 𝛼 with respect to the lens axis.
To keep the geometry compact, the scan angle 𝛽 has beenchosen to be the same as the focal angle 𝛼. Following [8],the ratio of the on-axis and off-axis focal length 𝑔 has been
determined from 𝑔 = 1/ cos(𝛼) as 1.0154. In [8], it has beenshown that, for any value of 𝑓 as defined in (1), if the focal
angle 𝛼 decreases, the value of 𝑔 approaches towards unity.
For a value of 𝑔 = 1.014 and a scan angle 𝛼 = 10∘ , 𝑓 has
been determined to be 0.5. In [8], it has also been shown thatthough the path length error will be at its minimum at 𝑓 = 1
for 15∘ < 𝛼 < 60∘ , however, for small angles such as 10∘ –15∘ ,the path length errors can be very small by using a value of 𝑓very close to 0.5. By keeping 𝑓 ≈ 0.5 a very compact lens canbe realized with a little compromise on negligible path length
error. To fabricate the lens using microfabrication techniques,
the lens geometry needs to be compact to optimize the yield.
Consequently, a ratio of 0.53 has been chosen for 𝑓 thatyielded a value for the off-axis focal length 𝐹 as 5.51 mm.
Accordingly, the on-axis focal length 𝐺 has been determined
to be 5.56 mm.
A high resistivity silicon substrate has been chosen tofabricate the lens as it has a very low loss tangent and absorption coefficient [7]. After initial design specifications were set,
the lens geometry was simulated in Rotman Lens Designer(RLD) software from Remcom Inc. RLD is a 2D solver thatuses the antenna aperture theory to determine electricallosses, lens geometry, microstrip tapers, and transmission
line lengths. The design parameters (geometric dimensions
and losses) were then optimized by using sidelobe levels,
phase errors, and sidewall curvature as the optimizing costfunctions. Following (2) below
𝑓𝑡 =

2
150
√
tan−1 𝜀𝑟 ,𝜋ℎ 𝜀𝑟 − 1(2)

where 𝑓𝑡 is the cutoff frequency and ℎ is the substrate thickness in millimeters, it was determined that a 200 micrometer
thick silicon substrate with a loss tangent of 0.0005 anddielectric constant of 11.8 would yield the optimum results
considering port coupling and sidelobe levels. A lower thickness substrate also prevents generation of the spurious surface
modes [10].
Considering the microstrip as a quasi-TEM line, theattenuation 𝛼𝑑 due to dielectric losses can be determinedfrom [6]𝛼𝑑 =𝑘0 𝜀𝑟 (𝜀𝑒 − 1) tan 𝛿𝑑Np/m,2√𝜀𝑒 (𝜀𝑟 − 1)(3)

where tan 𝛿𝑑 is the loss tangent. Therefore, a 1.5 micrometer
thick gold layer with conductivity 4.52 × 107 S/m was chosen

International Journal of Antennas and Propagation

3

Air (𝜀r = 1)

Silicon (𝜀r = 11.8)

Metal (gold)
1.5 𝜇m200 𝜇m(a)

Dummy ports

4

5

2

6

3

7

Array ports

15.6 mm
Beam ports

18

Dummy ports

Transmission lines
19.7 mm(b)

Figure 2: Rotman lens geometry. (a) Cross section. (b) Top view.

Table 1: Rotman Lens Design Specifications.
Parameter
Gold conductor thickness (𝜇m)

Value
1.5

Silicon substrate thickness (𝜇m)200

Silicon permittivity

11.8

Silicon loss tangent

0.0005

Length (mm) × width (mm)19.7 × 15.6

Center frequency
Bandwidth

77 GHz
8 GHz

Array factor (dB)

0

−10

−20

−30

0
Scan angle (deg)−50∘

Beamwidth
Focal angle

20
10∘

Scan angle
Beam ports
Array ports

10∘
3
5

Dummy ports
Element spacing

16
2.6 mmZ system

50 ohms

Port width

0.16 mm

Active beam 1
Active beam 2

50

Active beam 3

Figure 3: Array factor for three input ports excited separately.

to prevent degradation in the performance. The attenuation
𝛼𝑐 due to the conductor loss can be determined approximatelyfrom [6]𝛼𝑐 =𝑅𝑠Np/mm,(𝑍0 𝑊 )

(4)

4

International Journal of Antennas and Propagation
100
50

−9.5

Phase (deg)Amplitude (dB)

−9

−10

0
−50

−10.5

4

5

6
Array port number

7

8

−100

4

Beam port 3

Beam port 1
Beam port 2

5

6
Array port number

8

Beam port 3

Beam port 1
Beam port 2

(a)

7

(b)

Figure 4: (a) Beam to array coupling amplitude, (b) beam to array coupling phase.

−5

−40

−10
Amplitude (dB)

Amplitude (dB)

−15
−20
−25
−30
−35
−40
−45

−45
−50
−55
−60

2

4

6

Beam port 1
Beam port 28
1012Sidewall sample number

14

4

5

16

Beam port 3

6

7

8

Array port number
Beam port 1
Beam port 2

Beam port 3

Figure 6: Beam to array spillover coupling amplitude.

Figure 5: Beam to sidewall coupling amplitude.

simulation capabilities of XFDTD software were used to
optimize the final design specifications as listed in Table 1.
where 𝑅𝑠 = √𝜔𝜇0 /2𝜎 is the surface resistivity for conductor.
The loss associated with the connecting transmission lines
was calculated following 𝑒−𝛼𝐿 where 𝛼 = 𝛼𝑐 + 𝛼𝑑 [6]. The useof high resistivity silicon to reduce substrate loss and parasitic
components led to conduction between the semiconductor
wafer and the metal layers in a microstrip configuration. In
order to prevent current leakage to the substrate, a 100 nm
thick SiO2 layer was used between the high resistivity siliconwafer and the gold layers at the top and the bottom. A tapermatching technique with a taper width of 𝜆/2 was used to
avoid second-order mode. Through a simulation study in
RLD, a total of 16 dummy ports were determined to generate
low level sidelobes [8, 10, 11] while generating an optimumarray factor.
Figure 2(a) shows a cross-sectional view of the designed
Rotman lens and Figure 2(b) shows the top view of the lensin XFDTD 3D simulation environment. The 3D full wave

3. Simulation Results
The array factors of the designed lens for three input ports
excited separately are shown in Figure 3. From Figure 3, it isclear that, for each beam port excitation, the sidelobe levels
are less than −12 dB. Increasing the contour curvature as wellas number of the dummy ports can reduce the sidelobe levels.
Low curvature lens sidewalls have much higher sidelobes.
A Gaussian optimization method was used to determine
the number of dummy ports to match the determined lenscurvature. Phase error for individual beam port through all
the array ports is calculated as low as 0.5 degrees. This amountof error is very low as compared to 14.4∘ [4] and 24∘ [12] for
77 GHz lens.
Figure 4(a) shows beam to array coupling amplitude forindividual excitation of different beam ports. In worst case,

5

0

200

−5

100
Phase (deg)S-parameter (dB)

International Journal of Antennas and Propagation

−10

−100

−15
60

0

65

70

75

80

85

−200

90

6

6.5

7

Frequency (GHz)
SA7B1
SA8B1

SA4B1
SA5B1
SA6B1

8
7.5
Frequency (Hz)Array port 4

Array port 7

Array port 5

Array port 8

8.5

9
×1010

Array port 6
(a)

(b)

0

200

−5

100

Phase (deg)S-parameter (dB)Figure 7: 𝑆-parameter for beam port number one active, (a) magnitude, (b) phase.

−10

0
−100

−15
60

65

75

70

80

85

90

−200

6

6.5

Frequency (GHz)
SA7B2
SA8B2

SA4B2
SA5B2
SA6B2
(a)

7

8
7.5
Frequency (Hz)Array port 4
Array port 5
Array port 6

8.5

9
×1010

Array port 7
Array port 8(b)

Figure 8: 𝑆-parameter for beam port number two active, (a) magnitude, (b) phase.

array port number six has the most coupling amplitude withamount of −9.45 dB for beam port number one and three andalso −9.3 dB for beam port number two. Figure 4(b) shows
the amount of beam to array coupling phase.
The amount of beam to sidewall coupling amplitude(coupling of energy from the beam ports into the sidewalls)
among the beam ports and the 16 dummy ports (8 on each of
the side walls) is shown in Figure 5. About −10 dB maximum
coupling occurs for dummy ports that are close to array
ports as they will absorb most of the back scattered radiation.
Furthermore, beam ports to array ports spillover coupling
amplitude (coupling of energy from the beam ports to array
ports due to sidewall reflection) is between −58 dB to −43 dB
as shown in Figure 6.
Figure 7(a) shows the 𝑆-parameter amplitude andFigure 7(b) shows the 𝑆-parameter phase for beam port 1
active for a frequency range of 60 GHz to 90. The figures
show that 𝑆-parameter amplitudes for beam port number
one along the array ports number 4, 5, 6, 7, and 8 are −11.1 dB,−11.7 dB, −11.6 dB, −12.7 dB, and −10.9 dB, respectively, andthe associated phases are 3.76∘ , 47.44∘ , 87.91∘ , 130.40∘ , and
169.32∘ , respectively. Beam port number three has almost thesame results.
Figures 8(a) and 8(b) show the 𝑆-parameter values,
amplitude and phase, for beam port number two active.
At 77 GHz all the array port phases are around 63∘ with avariation of 2∘ .
Figure 9 shows the 𝑆-parameters (magnitude and phase)
for all the beam ports and array ports superimposed.
Figure 10 presents the return losses for individual excitation of the beam ports 1 to 3. From the figure, it is clear that thereturn losses associated with individual beam port excitation
are in the vicinity of −10 dB.
Crosstalk is one of the consequential factors when asignal is transmitted on transmission line and produces
an undesired effect in adjacent channels or transmission
lines. Crosstalk is usually created by unwanted capacitive,
inductive, or conductive coupling between transmission lines6

International Journal of Antennas and Propagation
0

S-parameter (dB)S-parameter (dB)

0

−5

−10

−5
−10
−15
−20

4

5

6
Array port number

7

60

8

70

75
80
Frequency (GHz)

S11
S22

Active port 3

Active port 1
Active port 2

65

85

90

S33

Figure 10: Return loss (dB) over frequency.

(a)

200
0
S-parameter (dB)

Phase (deg)

100
0
−100
−200

4

5

6
Array port number

7

8

−20

−30
60

65

70

75
80
Frequency (GHz)

SB2B1
SB3B1
SA5A4

Active port 3

Active port 1
Active port 2

−10

(b)

Figure 9: 𝑆-parameters for beam ports number one, two, and three
active, (a) magnitude, (b) phase.

since frequency is very high and size of the wafer is shrunk.
Figure 11 shows that incidence of crosstalk for all input andoutput lines is less than −12 dB.
Figure 12 shows that the insertion losses of beam ports
1 and 3, respectively, are 4.16 dB and for the beam port
2, the insertion loss is 3.29 dB at the center frequency of
77 GHz. It is to be noted here that, due to the maximum meshsize limitations associated with XFDTD, the 100 nm thicksilicon dioxide layer was not included in XFDTD simulation
to determine to insertion loss. If the 100 nm thick silicon
dioxide layer could be included in simulation, the insertion
loss would be much lower.
The lens efficiency can be defined as the ratio of the sumof the output power at all the array ports to the power fed at
one of the beam ports [4] and can be determined from the 3D
simulation results of the following transmission coefficients:
𝑃
Efficiency = out
𝑃in󵄨 󵄨2= ∑ 󵄨󵄨󵄨𝑆𝑛1 󵄨󵄨󵄨 = (𝑆41 2 + 𝑆51 2 + 𝑆61 2 + 𝑆71 2 + 𝑆81 2 ) .
𝑛

(5)

85

90

SA6A5
SA7A6

Figure 11: Crosstalk (dB) over frequency.

Figure 13 shows the lens efficiency as a function of
frequency for different beam port excitation. The efficiencyof beam ports 1 and 3 is 34.57%, and for beam port 2 and theefficiency is 46.8% at the center frequency of 77 GHz. This
proves that beam port number 2 which is in the center of thebeam ports can transmit higher power to the array ports.
Figure 14 shows the E-field pattern of current propagating
through the lens with beam port 2 active. The wave front willreach at all the array ports and only two closest of the dummy
ports with similar amplitude levels; therefore, most of thecurrent will be collected by the array ports. Table 2 compares
the key performance parameters of the designed Rotman lenswith two other similar microwave lenses.

4. Lens Fabrication Process
A 200 𝜇m thick high resistivity (>10,000 Ω-cm) silicon wafer
was RCA cleaned and a 100 nm thick layer of low temperatureoxide (LTO) was grown on both sides of the wafer. A 1.5 𝜇m
thick layer of gold was deposited by e-beam evaporation onto
the backside of wafer, Figure 15(a). LOR 30B lift-off resist
was then spin deposited at 2000 rpm and baked at 170∘ C
for 10 min, Figure 15(b). Shipley 1827 photoresist was thenInternational Journal of Antennas and Propagation

7

Table 2: Comparison of key features.
Parameter
Frequency (GHz)
Material
Thickness (𝜇m)Length × Width (mm)No. of beamports
No. of array ports
Phase error (deg.)
Insertion loss (dB)
Return loss (dB)Max. Efficiency (%)

Reference [13]24
RO3003 (𝜀𝑟 = 3)
508
7580
5
7
∗∗∗
1.85
Almost 10
32.30

Reference [14]60
Silicon and SiO2 (𝜀𝑟 = 11.7 and 3.9)
300 and 1.5
19.6 × 20
5
7
14.4
2
Almost 10
50

This paper77
𝜀𝑟 = 11.8
200
19.7 × 15.6
3
5
0.5
3.29
Almost 10
46.8

Insertion loss (dB)

10
8
6
4
2
0
60

65

70

75
80
Frequency (GHz)85

90

Beam port 1 active
Beam port 2 active
Beam port 3 active

Figure 14: E-Field simulation and current propagating for beam
port number two active.

Figure 12: Insertion loss (dB) over frequency.

Following DI water washing in the spin rinse dryer, a
1.5 micrometer thick layer of gold was deposited by e-beam
evaporation, Figure 15(e), on the topside of the processedwafer. The resist was then dissolved in Remover PG during anovernight soak to complete the metal lift-off step, Figure 15(f).
The wafer was then bonded to a carrier wafer with a layer ofcrystal bond to provide mechanical support during dicing tosingulate the devices. Figure 16 shows the final fabricated lens.

100

Efficiency (%)

80
60
40
20
0
60

5. Testing
65

70

75
80
Frequency (GHz)85

90

Beam port number 1
Beam port number 2
Beam port number 3

Figure 13: Power efficiency of the proposed Rotman lens for three
beam ports.

spin deposited at 2500 rpm and baked 3 minutes at 113∘ C,
Figure 15(c). The LOR and Shipley 1827 layers were thenpatterned with exposing the resist in the Karl Suss MA6aligner with 540 mJ/cm2 and developed for 3 minutes in a
MF-319 dish, Figure 15(d).

The lens was tested in a 77 GHz radar setup. To enable
testing, the lens was mounted on a special fixture along with a
77 GHz microstrip antenna array fabricated on a Duroid 5880
substrate as shown in Figure 17.
The complete test setup is shown in Figure 18. A tunable
X band source was used as the RF source which was thenconnected to an 8 times multiplier unit to output a millimeterwave signal in the vicinity of 77 GHz. The output of the 8times multiplier was then fed to one of the beam ports of theRotman lens. A 77 GHz horn antenna was used to pick up theradiation from the antenna. The horn antenna was connected
to an Advantest R3271A Spectrum analyzer (not in the figure)using a downconverter harmonic mixer. A screen shot of thespectrum analyzer is shown in Figure 19 where the outputof the harmonic mixer was superimposed on the original8

International Journal of Antennas and Propagation

100 nm silicon dioxide
200 𝜇mLOR resist

Silicon

1.5 𝜇m of gold layer(a)Shipley 1827

(c)

(b)

LOR patterning(d)

1.5 𝜇m of gold layer

Resist removal

(e)

(f)

Figure 15: Fabrication steps of the proposed Rotman lens.

Figure 16: Fabricated Rotman lens.

Figure 17: Fabricated Rotman lens combined with a microstrip
array antenna.

Figure 18: Complete test setup.

6. Conclusions
transmitted signal at 77 GHz. The frequency of the received
signal clearly indicates that the Rotman lens is functioning
properly as an integral part of the radar unit.

The design, fabrication, and testing of a high resistivity
silicon wafer based 77 GHz microstrip Rotman lens have beenpresented. Full wave simulation of the lens was carried out

International Journal of Antennas and Propagation

Figure 19: A screen shot of the spectrum analyzer showing both thetransmitted and the received signal to verify the functionality of theRotman lens.

using XFDTD software and the simulation results show thatthe lens exhibits very low phase error, compact size, andexcellent RF performance. The lens has been fabricated using
a lift-off process using a silicon dioxide layer between the top
metal layers the silicon substrate to reduce substrate loss.
A functional test of the lens has been carried out using
a 77 GHz radar setup and the test results show that thelens is working fine as designed. Further testing of the lensis in progress. The relatively less expensive microfabricatedRotman lens can be used to realize a high performance low
cost MEMS based radar sensor for automotive applications
by replacing expensive analog and/or digital beamformers as
currently being used in commercially available automotive
radar sensors. The developed Rotman lens can easily be
integrated with the Radar front end circuitry and silicon
based microstrip antenna array to pave the way to realize low
cost high performance radar on a chip.

Conflict of Interests
The authors declare that there is no conflict of interests
regarding the publication of this paper.

Acknowledgments
This research has been supported by the Natural Sciences andEngineering Research Council of Canada (NSERC)’s Discovery Grant no. RGPIN 293218, Ontario Centres of Excellence(OCE)’s Grant no. IC50659, and Auto21 NCE Canada’s Grant
no. F503-FSS. The authors would like to greatly acknowledge
the packaging and testing support provided by Advotech Inc.
of Tempe, Arizona. The authors also would like to acknowledge the supports provided by Canadian Microelectronics
Corporation (CMC Microsystems).

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Journal of

Journal of

Electrical and Computer
Engineering

Robotics
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Volume 2014

Volume 2014

VLSI Design
Advances in
OptoElectronics

International Journal of

Navigation and
Observation
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Volume 2014

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Chemical Engineering

Antennas andPropagation

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Volume 2014

Aerospace
Engineering

Volume 2014

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Volume 2014

Hindawi Publishing Corporation
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Volume 2014

Volume 2014

International Journal ofInternational Journal ofInternational Journal of

Modelling &Simulation
in Engineering

Volume 2014

Active and Passive
Electronic Components

Shock and Vibration

Acoustics and Vibration

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Hindawi Publishing Corporation
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Hindawi Publishing Corporation
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Volume 2014

Volume 2014

Advances inVolume 2014