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7 August 2019 Laser scanner module with large sending aperture and inherent high angular position accuracy for three-dimensional light detecting and ranging
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Abstract

One design of the state-of-the-art laser scanner systems in automotive applications is based on oscillating mirror modules. The requirement of a large mirror surface for eye-safe transmission beams and long measurement distances is a major drawback for fast and reproducible scanning. Tolerances of angular positioning, position sensing, and vibrational perturbations limit the position accuracy of such a mirror and, thus, the accuracy of the transmission spot position in the field of view (FoV). Our approach for a scanner module with maximum transmission beam diameter combines a microlens array with an objective lens for generating one optical telescope assembly for each angular scan position exclusively. Aperture stops define the beam positions in the FoV and avoid positioning errors caused by angle deviations of the scanner mirror. This increases the reliability of the angular position accuracy of the scanner module significantly. To minimize the shadings between adjacent scan spots in the target distance, created by beam cutoffs at the aperture stop of the objective lens, an array of optimized microwedge prisms is provided in combination with the microlens array. Therefore, we can increase the throughput of transmission power into the FoV and improve the measurement distance, especially at large scan angles.

1.

Introduction

Typically, there are two system principles of time-of-flight distance measurement, namely, scanner systems and three-dimensional camera/flash light detecting and ranging (LiDAR) systems. While scanner systems illuminate and collect each object point in the field of view (FoV) sequentially,1,2 flash LiDAR systems measure the entire FoV simultaneously in one snapshot.3,4 Hence, high frame rates are realized without smearing effects of moving objects. Owing to a solid-state setup, each measurement point relates to a fixed field angle and can be detected by one certain receiver element. This allows a precise spatial assignment of object points in the FoV, which leads to a high absolute angular accuracy of the system. The maximum measurable range of flash LiDAR systems is smaller with respect to scanner modules if the FoV size and transmission power per pulse are identical. The advantage of scanner modules regarding measurement ranges results from a higher level of receiving signal due to the concentration of the laser power to one measurement point, while flash LiDAR requires simultaneous illumination of the entire FoV. In terms of realizing the specifications of obstacle recognition and FoV surveillance in automotive applications, typically, scanner systems are the preferred technology in the market.5,6 Concerning transmission power consumption, the trade-off between meaningful measurement distances (<100  m)6,7 and horizontal FoV dimensions (>±20deg)8 is obligatory for flash LiDAR systems. The power consumption exceeds by factors compared to that of scanner systems, which is why the common operating range of flash systems is around 0.1 to 25 m.3,7

The state-of-the-art laser scanner systems in automotive applications are based on transmission signal deflection modules with large, heavy and expensive oscillating or rotating mirrors. For transmitting eye-safe laser beams and collecting maximum amounts of scattered and reflected signal of measured targets in the FoV, large sending and receiving apertures, which means large scanning mirror surfaces, are essential. The reachable distance up to 200 m and the corresponding distance uncertainty is typically in the range of 2 mm to 10 cm, depending on the environmental conditions and ambient illumination. The range of angular resolution is around 0.01 deg to 0.5 deg with an angular uncertainty up to 0.015 deg.59 However, the large construction size, the fine mechanics setup, as well as low potential of automated high volume manufacturing seem to be the major drawbacks of this kind of scanner design.

Replacing the fine-mechanical deflection mirror by a microelectromechanical system (MEMS) scanning mirror overcomes the large construction size. However, a smaller scanning mirror size reduces the sensor aperture and, therefore, the measurement distance. Consequently, a mirror array or a very large single MEMS mirror setup is in the range of typically 1- to 11-mm diameter.912 On the one hand, the cost factor rises directly with the active mirror area.13 On the other hand, the requirement of a large MEMS mirror surface is a major drawback for fast and reproducible scanning. Tolerances of angular positioning, position sensing, as well as vibrational perturbations, e.g., externally induced shock lead to mirror-deflection noise, which is projected directly into the FoV. Thus, the resulting position noise of measurement spots in the FoV is the main reason for the limitation of the accuracy of spot positions.

Add-on closed-loop feedback circuits realize the measurement of the actual mirror position. The position can be readout either optically on the basis of the position of reflected signal on a monitor diode at the backside of the mirror14 or piezoelectrically through the measurement of torsion at the mirror springs while oscillating.15 However, with this technique several vibration modes can cause ambiguous position feedback. A further possibility is to measure the capacity variation depending on the overlapping depth of finger structures at the mirror.16,17 The reachable angular uncertainty in closed-loop circuit modules is located in the range around 0.015 deg to 0.1 deg.

By increasing the mirror surface, the inherent dynamic deformation δμM increases with the mirror diameter dμM (δμMdμM5) while oscillating, especially in the turning points. In addition, the static deformation induced by tensions between particular mirror layers becomes more relevant.18,19 To scale-up the mirror thickness tμM (δμM1/tμM2) is one possibility of reducing the deformations δμM. This again leads to higher mass of the mirror mμM and, respectively, higher inertia IμM (IμMdμM2, assumed a round mirror surface). Higher inertia and higher rotating angles φμM correspond to increasing torsional stress κS in the mirror springs (κSIμMφμM). For fast retraction, a higher stiffness of springs is advantageous, which increases the stress level and the postoscillation. Complementarily, this reduces the shock and vibration resistivity.1921 In conclusion, regarding lifetime and robustness aspects, the target for the operation of a scanning mirror module is minimum dimensions dμM and low deflection angles φμM.

As presented in Refs. 22 and 23, scanner systems, including micro-optics, are an alternative approach for maximization of the sending and receiving apertures without increasing the mirror surface. Here, the assembly of one micro-optical element with an objective lens generates an optical telescope for each separate scan position exclusively.2426 The maximum transmission beam diameter and the maximum receiving aperture of these systems accordingly depend on the combination of microlens and objective lens. Transmission beam cutoffs at the aperture stop of the objective lens turn out to be an inherent system drawback while scanning over a microlens array. The consequence is a significant loss of measurement distance and an intermittently sampled FoV at large field angles.

In this paper, we present the approach for a scanner module based on prior work.22,23 This scanner module improves the position accuracy of measurement spots in the FoV without any shadings in between. Among others, we show in detailed simulation and measurement the influence of spot position noise and beam cutoffs on the transmission power. Usually, mirror-deflection noise is the dominant limitation for the position accuracy of the measurement spots. We design the dimensions of aperture stops between the microlenses and the transmission beam diameter on the microlens array carefully in order to suppress the direct projection of the angular noise of the rotating mirror into the FoV. Therefore, our new design approach leads to a system with inherent high measurement spot-position accuracy. In addition, the design of a specific microwedge prism for each particular microlens avoids previously described transmission beam cutoffs at the aperture stop of the objective lens. This avoids resulting shadings between adjacent scan spots in the FoV and increases the throughput of transmission power into the FoV. The consequence is the improvement of measurement distance, especially at large scan angles.

2.

Laser Scanner Module

The geometrical correlations of the actual scanning principle are described in Sec. 2. We also show the system limitations regarding reachable transmission beam diameters and the correlated maximum allowed transmission power. This consideration is in conformity to the IEC 60825-1:2014 standard.

2.1.

Principle of the Laser Scanner Module

In this paper, we present a laser scanner system (Fig. 1), which realizes large transmission beam diameters with minimum mirror dimensions and rotation angles. We derive an equation for the two-dimensional (2-D) rotating angles φμM of the oscillating mirror μM to hit the center position of each microlens on the microlens array MLAT with the deflected transmission beam optimally [see Fig. 2(a)]. The necessary angle depends on the incoming transmitter vector VTin on mirror μM, the inclination angle βMT of mirror MT, and the center positions of the microlenses on the microlens array MLAT. For better clarity, Figs. 1(a) and 2 just show the yz-plane. The rotation angle of the mirror μM around the y-axis with φμM,y in the xy-plane is derived analog to the rotation angle around the x-axis with φμM,x in the yz-plane. We also show the derivation of the divergence angle θT in dependence on the focal length fμL and the diameter dμL of the microlens. It also depends on the distance aμL of the microlens to the objective lens (Lens LO) and its focal length fLO.

Fig. 1

Setup of the laser scanner module. For better clarity, just the yz-plane is shown here. (a) Schematic depiction of the optical path inside the scanner module, with the help of two sequentially transmitted laser beams (T0 and Tn for one mirror position each) and the corresponding received signal on the detector after scattering at the target (R0 and Rn). (b) Demonstrator setup with (1) Crylas FDSS 532-Q laser, (2) Thorlabs CPS532-C2 alignment laser, (3) Smaract STT-2013 2-D motorized optical mount, (4) Olympus M. Zuiko Digital ED 25 mm F1.2 Pro standard objective lens, and (5) IDS UI-1450-C USB 2.0 camera (for testing issues).

OE_58_8_087101_f001.png

Fig. 2

Geometrical operation principle of the laser scanner module. For better clarity, just the yz-plane is shown here. (a) Scanner section with the geometrical correlation of the rotational angle φμM,x of the rotating mirror μM depending on the transmission vector VTRes,n. This position vector describes the vector from the rotating position of mirror μM to the center position of each particular microlens on the microlens array MLAT. (b) Simplified beam-shaping section with correlation between the divergence angle θT and the z-positioning aμL of the microlens array MLAT to the objective lens LO.

OE_58_8_087101_f002.png

The scanner module provides the classification of laser class 1 in conformity with the IEC 60825-1:2014 standard. Our system incorporates an optimized micro-optical array for noise suppression of the transmitted spot position in the FoV. The optics solution additionally avoids shadings in the FoV for high light-collecting efficiency even at large field angles. These characteristics allow high robustness against vibrational perturbations and front lens contaminations, as well as constant angular resolution within maximum measurement range. The separation of optical paths in the setup gives a high stray light insensitivity of transmission signal reflections at components in the transmitter path. In addition, the setup allows separate optical design of the microlens arrays for particular transmission and receiving requirements.

Figure 1(a) shows the optical operating principle of the scanner module and, respectively, the equivalent demonstrator setup [see Fig. 1(b)]. Outgoing from the laser source (transmitter), the transmission beam (T0 and Tn for one exemplarily angular mirror position each) is deflected via a MEMS mirror μM to the lens LT, which converts the rotational movement of mirror μM to a parallel shift to the optical axis. This again realizes a sampling of each microlens perpendicular to its principal plane on the microlens array MLAT. Depending on the focal length fLT of lens LT, the necessary rotation angle of mirror μM correlates directly to the resulting parallel shift of T0 to Tn. The beam-shaping section, consisting of one microlens and the objective lens LO, generates an optical telescope assembly for each scan position and, respectively, spot position in the FoV. Therefore, the amount of microlenses on the microlens array MLAT is equal to the amount of spot positions in the FoV, assuming no subsampling on the detector. Concerning eye-safety requirements, each microlens widens the transmission beam to the size of a less critical diameter dTLO at the system exit plane of the objective lens LO [see Fig. 2(b)]. Despite a beam widening, the advantage compared to a conventional telescope setup is no loss of field angle in the FoV at all. After scattering at the target, the receiving signal (R0 and Rn) reaches a separate receiver path via the objective lens LO and the beam splitter. Each angular mirror position of the mirror μM corresponds to one particular incidence angle of the receiving signal (see R0 and Rn). Accordingly, the deflection of the receiving signal can be compensated to one static position on the detector with the advantage of minimum necessary detector dimensions. Such a system setup is favorable for the design of a microlens array MLAT in the transmission path with additional aperture stops between each microlens. This design suppresses spot position noise in the FoV caused by the limited angular accuracy of the mirror μM. However, the microlens array MLAR in the receiving path can be designed with an optimized diameter of each microlens without additional aperture stops for maximum light-collecting efficiency.

The relation of the necessary 2-D rotation angles φμM,x and φμM,y for deflecting the transmission beam depending on the corresponding center position of each sampled microlens is illustrated in Fig. 2(a). For the following calculation, the rotation point of the rotating mirror μM is defined as the point of origin of the coordinate system (marked with the red point).

The vector VTRes,n describes the position vector from the rotating point of mirror μM to the center position of each single microlens. In consequence, this vector is defined with the given coordinates of the microlens array MLAT (xMLAT,0, yMLAT,0, and zMLAT) as well as the pitch between each microlens PitchMLAT and the number nμL of each particular microlens in x- and y-directions

Eq. (1)

VTRes,n=(nμL,x·PitchMLAT,x+xMLAT,0nμL,y·PitchMLAT,y+yMLAT,0zMLAT).

The vector VT,n with VT,n=VTRes,nVTLM,nVTμL,n describes the transmission beam from the rotating point of mirror μM to lens LT to hit finally the center position of each microlens. The vector VTLM,n describes the transmission beam after refraction on lens LT and VTμL,n is the vector after deflection at the mirror MT. The inclination angle βMT of mirror MT, the angle of lens LT φT0,x=2·βMT, and its focal length fLT are given by the system design as well. Hence, VT,n can be expressed in dependence on these known system design parameters as

Eq. (2)

VT,n=[xμL,nsin(φT0,x)·(fLTC1)·sin(φT0,x2)·1tan(φT0,x)cos(φT0,x)·(fLT+C1)·sin(φT0,x2)·tan(φT0,x)],
where xμL,n=nμL,x·PitchMLAT,x, yμL,n=nμL,y·PitchMLAT,y, zμL,n=yμL,ntan(βMT), C1=zμL,n2+yμL,n2, and zμL,n=yμL,ntan(βMT).

Now, we can determine the necessary rotation angle φμM of mirror μM to deflect the known incoming transmitter vector VTin to the vector VT,n. In a first step we have to calculate the normalized normal vector to the mirror surface μM Vu,N=Vu,Tin+Vu,T,n, where Vu,Tin is the given normalized vector of the laser source and Vu,T,n is the known normalized vector of the transmission beam after deflection on mirror μM. Then, we calculate VRot with the inverse direction cosine matrix VRot=DCMy1·Vu,N, which describes the rotated vector Vu,N to the optical axis in z-direction. The 2-D rotation φμM to deflect the transmission beam to the center position of each microlens finally can be described as

Eq. (3)

φμM=  (φμM,xφμM,y)=[tan1(VRot,yVRot,z)tan1(Vu,N,xVu,N,z)].

In Fig. 2(b), the beam-shaping section of the scanner module is depicted in simplified form with just one microlens. It shows the correlation of the distance aμL in z-direction between one microlens and the objective lens LO. The resulting divergence angle θT, with help of the paraxial approximation, is expressed as

Eq. (4)

θT=tan1[dμL2·fμLaμL·dμL2·fLO·fμL1+(dμL2·fμL)2·aμLfLO],
where dμL and fμL are the diameter and the focal length of one microlens, respectively. The parameter fLO describes the focal length of the objective lens LO. Hence, the divergence angle can be adjusted by varying the distance aμL between the microlens array MLAT, and the objective lens LO. In principle, correct adjustment of θT offers the capability to sample the FoV without any gaps between adjacent scan spots in the FoV.

2.2.

Limitations of the Transmission Beam Diameter and Transmission Power

The geometrical correlation between reachable FoV dimensions and angular resolution αTr, coupled with large transmission beam diameter dTLO, is one fundamental limitation of scanner systems including micro-optics for transmission beam widening. The theoretically feasible transmission beam diameter on the principal plane of the objective lens LO is limited by its diameter dLO. We assume planoconvex spherical lenses with a refraction index nMLAT of 1.5 and a maximum radius of curvature RμL of half the diameter dμL of one microlens on the microlens array MLAT. In consequence the lower limitation of focal length of one microlens is fμL,min=2·RμL, which results in a f-number F#μL=fμL/dμL1.27 Here, a microlens with F#μL=1 defines the minimum possible focal length fμL,min and, thus, a maximum beam widening of the transmission beam. In this scanner system, the number of scan spots in the FoV is equal to the number of microlenses on the microlens array MLAT, with no subsampling on the detector assumed. Consequently, the pitch of the microlens array MLAT corresponds to the projected PitchTr between the measurement spots in the FoV as well. This again is related to the system angular resolution αTr=tan1(dTr/zTr), with the transmission spot diameter on the target dTr at the target distance zTr [see Fig. 2(b)]. For the following calculations, the PitchMLAT of the microlens array MLAT equates to the diameter of one microlens dμL. Outgoing from a chosen focal length fLO of the objective lens LO and the resulting horizontal FoV FoVh, the system-required pitch of the microlens array MLAT is given as

Eq. (5)

PitchMLAT=dLO·tan(αTr)2·tan(FoVh2).

This is illustrated in Fig. 3(a) with the example of one standard objective lens LO1 with a FoVh of 47 deg and two wide-angle objective lenses LO2 and LO3 with 57 deg and 84 deg, respectively. The necessary focal length fμL of the microlens with

Eq. (6)

fμL=dLO2·tan(αTr)4·tan(FoVh2)2·dTLO,
to generate a defined transmission beam diameter dTLO on the principle plane of the objective lens LO can be taken from Fig. 3(b). With the specified angular resolution of αTr=0.286  deg, which is based on a spot diameter in the FoV dTr=0.5  m at zTr=100-m target distance, only objective lens LO1 and LO2 can meet the requirements. This result is with respect to the minimum reachable f-number of the microlens on the microlens array MLAT. As shown in Fig. 3(b), with the example of LO3, the necessary focal length fμL (dashed green line) is significantly lower than the theoretically feasible minimum focal length fμL,min (dotted green line). The result of Fig. 3(b) shows the limitation of a maximum FoV of the defined system of ±28.5  deg horizontally with fμL=99.6  μm and PitchMLAT=99.6  μm at dTLO=20  mm (red dashed and dotted line one above the other).

Fig. 3

Geometrical correlation between microlens parameters, horizontal FoV FoVh, and transmission beam diameter dTLO with predefined angular resolution αTr=0.286  deg (for three objective lenses LO1, LO2, and LO3 with a FoVh of 47 deg, 57 deg, and 84 deg). (a) The resulting PitchMLAT to reach the specified resolution with given objective lenses and their focal lengths fLO. (b) The resulting focal length fμL and its lower limit fμL,min of one microlens to realize a specific transmission beam diameter dTLO, exemplarily for dTLO=20  mm. This is based on the resulting PitchMLAT from (a), which is given by the defined objective lenses LO1, LO2, and LO3.

OE_58_8_087101_f003.png

Theoretically, our scanner concept realizes transmission beam diameters up to 20 mm over the entire FoV, limited by the clear aperture at the system exit plane on the objective lens LO. Because of the micro-optical-based beam widening, the transmission beam can be deflected with a dimensionally small oscillating mirror. In conformity with the IEC 60825-1:2014 standard, our scanner system allows, with an assumed spot diameter of 20 mm, maximum permissible exposures (MPEs) of <4  kW peak power within 2.5-ns full width at half maximum (FWHM) at a wavelength of 900 nm. A high level of transmission power is indispensable for long measurement distances, considering a minimum measurable value of the receiving signal PR in a range of 109 to 1012  W at the detector. This is based on a huge loss of backscattered signal power even at ideal 100% scattering targets in the FoV with respect to the transmission signal PT (e.g., PRPT·108 at zTr=100  m). We can express the behavior of the receiving signal depending on the ratio of etendues28 GLO to GTr of the receiving aperture and that of the backscattered signal at the target (PRGLOGTr). The MPE power29,30 PT,max scales with the beam diameter dTLO with PT,maxdTLO2.

For our experimental setup we choose a light-sensitive objective lens LO with a f-number of 1.2 and a FoVh of ±23.5  deg (necessary PitchMLAT=124.8  μm) (see Table 1). The microlenses used for the transmission beam widening allow a transmission beam diameter dTLO of 10.41 mm and consequently a possible transmission pulse-power maximum PT,max of up to 1 kW, consistent with the IEC 60825-1:2014 standard.

Table 1

System specifications of the experimental setup of the laser scanner module.

SymbolValueDescription
FoVh±23.5  degHorizontal FoV
zTr0.9 mTarget/object distance
dTLO10.41 mmTransmission beam diameter at objective lens LO
PT<1  kWTransmission pulse power at 900-nm wavelength
tPulse2.5 nsTransmission pulse length

3.

Suppression of Spot Position Noise

In Sec. 3, we show the geometrical correlation between the mirror-deflection noise of the oscillating mirror and the position accuracy of measurement spots in the FoV. We describe how to suppress this position noise in the FoV with the approach of a careful design of the aperture stop dimensions between the microlenses and the transmission beam diameter on the microlens array. Outgoing from the mirror-deflection noise, we show the effect on varying transmission power and spot homogeneity. Further, we give an estimation of the transmission power consumption for a position-noise-suppressed system.

3.1.

Correlation between Mirror-Deflection Angle and Spot Position in the Field of View

One deflection angle of the oscillating mirror μM correlates with one particular position of the transmission beam on the microlens array MLAT in the xy-plane (see Fig. 4). Outgoing from this position, the transmission beam is projected through the objective lens LO into the FoV and directly defines the field angle wFoV and spot position on the target. Consequently, angular mirror-deflection noise of the oscillating mirror leads directly to spot position noise on the microlens array and finally to spot position noise in the FoV. Next to externally induced perturbations, mirror-deflection noise is the dominant limitation for the position accuracy of measurement spots.

Fig. 4

Scanner section with the geometrical correlation of the resulting spot position on the microlens array and in the FoV in the xy-plane depending on the mirror-deflection angle φμM,x of the rotating mirror μM. These positions are given with the radial distance hT,y and hFoV,y from the optical axis of the objective lens LO. For better clarity, just the yz-plane is shown here.

OE_58_8_087101_f004.png

On dependence of an arbitrary rotation angle φμM, we can express the functional correlation to the corresponding spot positions in the xy-plane on the microlens array MLAT and in the FoV. Through the DCM rotation convention with Vu,N=DCMy[DCMx·Vu,z], in a first step, we calculate the normalized normal vector of the mirror surface Vu,N, where Vu,z is the unit vector in z-direction. With the normalized transmission beam vector

Eq. (7)

Vu,T=Vu,Tin+2·(Vu,N·Vu,Tin)·Vu,N,
the angle φT,T0=cos1(Vu,T·Vu,T0) in the plane between Vu,T and normalized vector VT0 on the optical axis of lens LT, the deflected transmission beam vector VT is given as

Eq. (8)

VT=Vu,T·fLTcos(φT,T0).
Here, Vu,Tin describes the known normalized vector of the laser source and fLT denotes the focal length of lens LT. The resulting spot position of the transmission beam in the xy-plane on the microlens array MLAT and in the FoV is expressed by its radial distance hT and hFoV, with respect to the optical axis of the objective lens LO, as

Eq. (9)

hT=  [VT,x(VT0,yVT,ycos(φT0,x))]
and

Eq. (10)

hFoV=(zTrfLO)·(hT,xfLOhT,yfLO).

The angle between the lens LT and the optical axis of the system is given as φT0,x. The variable zTr represents the target distance and fLO is the focal length of the objective lens LO. The side inversion because of the reflection on mirror MT and the projection through the lens LO are considered with a negative sign each.

With Eq. (9) we can derive the propagation of the angular error DφμM of the oscillating mirror μM, which is based on the mirror-deflection noise, to the resulting position error DhT on the microlens array in the xy-plane, as

Eq. (11)

DhT,y=|hT,y(φμM,x,φμM,y)φμM,x|·DφμM,x+|hT,y(φμM,x,φμM,y)φμM,y|·DφμM,y.

Here, exemplarily the y-component is shown (see Fig. 4). We assume an equal probability density distribution of this mirror-deflection noise for calculating the angular uncertainty of the mirror μM with uφμM=DφμM,x/3. Accordingly, the combined uncertainty of the transmission beam position on the microlens array MLAT for two uncorrelated input parameters is expressed as

Eq. (12)

uhT,y=(C2)2·uφμM,x2+(C3)2·uφμM,y2,
where C2=hT,y(φμM,x,φμM,y)φμM,x and C3=hT,y(φμM,x,φμM,y)φμM,y.

Based on Eq. (10), the position error in the FoV DhFoV,y and the spot position uncertainty in the FoV uhFoV,y are derived analog to this.

3.2.

Approach for Suppression of Spot Position Noise with Aperture Stops

In a scanner system with a microlens array in the beam-shaping section with high optical fill factor as shown in Figs. 5(a) and 6(a), the spot position noise on the microlens array MLAT is directly projected into the FoV. Here, a position error of the transmission beam T0,DhT from the reference position T0 on the microlens array MLAT, leads to a position error DhFoV in the FoV. In our approach, we suppress this direct projection of spot position noise on the microlens array into the FoV with aperture stops between the microlenses and a reduction of the optical fill factor [see Fig. 5(b)].

Fig. 5

Schematic principle of the effect of the mirror-deflection noise on the measurement position in the FoV. (a) Transmission path with a microlens array without additional aperture stops; here MLAT with dμL=PitchMLAT=dT,MLAT. (b) Transmission path with a microlens array with additional aperture stops; here, MLAT,NS with 2·dμL=PitchMLAT,NS=dT,MLAT for suppression of spot position noise in the FoV.

OE_58_8_087101_f005.png

Fig. 6

Microlens arrays utilized in the experiments. (a) MLAT with PitchMLAT=dμL (Zeiss SmartZoom m5, 34× Zoom, Plan Apo D 1.6×/0.1 FWD 36 mm, MLA PowerPhotonics PP-FRF-1464). (b) MLAT,NS with PitchMLAT,NS=2·dμL (Zeiss SmartZoom m5, 34× Zoom, Plan Apo D 1.6×/0.1 FWD 36 mm, MLA PowerPhotonics PP-FRF-1465). (c) Transmission beam directed on the center of a microlens (1), and on the transition zones between the microlenses (2) and (3), with HRS015B stabilized HeNe laser.

OE_58_8_087101_f006.png

The diameter of these aperture stops equates to the diameter of the microlenses dμL. We design the center-to-center distance of adjacent microlenses PitchMLAT,NS to twice the value of the maximum position error DhT,y of the transmission beam on the microlens array and coat it with light impermeable material

Eq. (13)

PitchMLAT,NS=dT,MLAT=2·(dμL2+DhT,y).

Considering this range of the spot position noise, we have to increase the transmission spot diameter dT,MLAT on the microlens array to the same extent, to ensure fully illuminated microlenses at any time.

Hence, the range of suppression of spot position noise depends on the aperture stop dimension, the pitch between adjacent microlenses and the transmission spot diameter pointing on the microlens array MLAT,NS. Our microlens design offers a pitch of twice the diameter of the microlens PitchMLAT,NS=2·dμL [see Fig. 6(b)].

With such a microlens array and a focal length fLT=75  mm of lens LT [see Figs. 1(a) and 2(a)], we achieve to suppress a spot position noise in the FoV with a position error of DhFoV0.25  m (Dht62.5  μm on MLAT,NS). This means a suppression of mirror-deflection noise with an angular uncertainty of uφμM,x=0.013  deg. For small deflection angles, the spot position uncertainty can be assumed as constant. In a system without a position-noise-suppressing microlens array and with an angular resolution αTr=0.286  deg and a FoVh=±23.5  deg (see Sec. 2.2), this spot position error DhFoV equates to half the target diameter dTr, which is to be resolved. This means a spatial degradation of angular resolution up to αTract=½·αTr. In worst case, targets within the diameter dTr can be missed, because of emerging gaps between sequential scan spots.

3.3.

Effects of Spot Position Noise on Transmission Power and Spot Homogeneity

In addition to spot position noise in the FoV, a reduction of the actual transmission power and spot homogeneity is the consequence, if illuminating a microlens not centrically. This reduction depends on the dimension of spot position error of the transmission beam position on the microlens array to the center position of the microlens in the xy-plane. Variations of transmission power and spot homogeneity directly affect the maximum reachable measurement distance and eye safety.

To evaluate the effect of spot position noise on the transmission power, with a static setup we consciously displace the transmission spot on the microlens array to the center position of the microlens in the y-direction. In Figs. 7(a) and 7(b), we experimentally can show this reduction of transmission power. Outgoing from the reference spot T0 with hT,y=0, we displace the spot up to T0,hT with hT,y=½·dμL of the microlens diameter. For a microlens array without a position-noise-suppressing design [see Figs. 6(a) and 6(c)], eccentric illumination of the microlens means partial illumination of translucent transition zones between the microlenses on the array MLAT. We can measure a reduction of transmission power PT of about 25% from centric to eccentric illumination of the microlens.

Fig. 7

Transmission power PT on the target in the FoV, depending on the displacement hT,y (in y-direction) of the transmission beam on the microlens array MLAT. The illumination is shown for a centric (T0) position and an eccentric (T0,hT) position relative to the microlens position. The laser source is a Crylas FDSS 532-Q laser and the transition zones of the microlens array are translucent. (a) The correlation of the transmission spot position on the microlens and its projection into the FoV. (b)The resulting curve of transmission power PT of the measurement spot in the FoV at zTr=0.90  m, depending on axial displacement of the transmission beam from the center hT,y=0 to ½·dμL of the microlens diameter. (c) Cross section of the transmission spot intensity ISpot, in gray value GV in the FoV at zTr=0.90  m, T0: centric illumination of the microlens hT,y=0. (d) Cross section of the transmission spot intensity ISpot in gray value GV in the FoV at zTr=0.90  m, T0,hT: eccentric illumination of the microlens hT,y=½·dμL (8-bit IDS UI-1450-C USB 2.0 camera).

OE_58_8_087101_f007.png

To express the homogeneity HI with respect to the spot intensity ISpot, we use the mean absolute deviation of all intensity values along the spot cross section in the y-direction to the arithmetic mean value I¯Spot of the intensity. We just consider values inside the microlens aperture. Related to the measurement, the parameter i defines camera pixels in the range from imin to imax with ±½·dμL [see the red dashed lines in Figs. 7(c) and 7(d)]. With normalization of the mean deviation to I¯Spot, we can write the homogeneity as

Eq. (14)

HI=1(1Ni=iminimax|ISpot,iI¯Spot|)·1I¯Spot,
where N is the number of intensity summands. In Figs. 7(c) and 7(d), we show the corresponding cross section of the transmission spot intensity in the FoV for the reference and the displaced transmission beam T0 and T0,hT. A significant decrease of the spot homogeneity from HIT0=73.7% to HIT0,hT=51.1% can be observed with an increase in the displacement hT,y.

For a microlens array with a position-noise-suppressing design [see Fig. 6(b)], eccentric illumination of the microlens means a direct increase of the illuminated area on the light impermeable aperture stops. This results in a transmission beam that only partially transmits through the microlens array, which is shown simulatively in Fig. 8.

Fig. 8

Transmission spot through the microlens, depending on the displacement hT,y of the transmission beam on the microlens array MLAT,NS. The illumination is shown exemplarily with two ratios (A and B) of beam diameter dT,MLAT to microlens diameter dμL on a microlens array with position-noise-suppressing design. This is illustrated for a centric illumination of the microlens with hT,y=0 (upper figure) and an eccentric illumination with hT,y=½·dμL (lower figure). The figures on the right of (a) and (b) show the detailed view of one microlens. (a) A: dT,MLAT=dμL. (b) B: dT,MLAT=2·dμL.

OE_58_8_087101_f008.png

The consequence is a stronger reduction of transmission power PT in the FoV compared to the microlens design shown in Fig. 6(a). The transmission power PT in dependence on the displacement hT,y of the illuminated position to the center position of the microlens can be expressed as

Eq. (15)

PT=ISpot(x,y)·AμL[x,(yhT,y)]dydx.
Here, AμL is the translucent area of the microlens, which is convoluted with the spot intensity ISpot. If we assume a Gaussian intensity profile of the transmission beam, we can write the equation as

Eq. (16)

ISpot(x,y)=ISpot,0·12π·σ2·exp(x2+y2)2·σ2,
where σ is the standard deviation of the beam profile.31 Figures 9(a) and 9(b) show the spot intensity ISpot and the transmission power PT. This is illustrated for two ratios of beam diameter to microlens diameter, A: dT,MLAT=dμL and B: dT,MLAT=2·dμL with identical transmission power PTA=PTB. On this example, we evaluate the behavior of the spot homogeneity and the transmission power depending on the displacement hT,y. The spot diameter dT,MLAT is defined with ±2σ.

Fig. 9

Effects of spot position noise on transmission power and spot homogeneity. (a) Cross section ISpot in y-direction. (b) Transmission power PT transmitting through one microlens. With two ratios of beam diameter dT,MLAT to microlens diameter dμL,. A: dT,MLAT=dμL and B: dT,MLAT=2·dμL. This is shown in dependence on the displacement hT,y of the transmission beam from the center hT,y=0 to hT,y=½·dμL relative to one microlens, described with T0 and T0,hT.

OE_58_8_087101_f009.png

In Table 2, we show the summary of the simulative evaluation based on the graphs given in Fig. 9. With respect to the spot homogeneity HI and the relative loss of transmission power PTLoss, we can see the advantageous effect of a diameter ratio B. One negative effect of position-noise-suppressing design of the microlens array is a significant increase of the necessary transmission power compared to a microlens design shown in Fig. 6(a). Thus, the light loss at the aperture leads to a necessary increase of transmission power about the factor 2.19 for a design illustrated in Fig. 8(b).

Table 2

Evaluation of the spot homogeneity HI and the loss of transmission power PTLoss. Here, these values are given in dependence on the transmission spot diameter dT,MLAT (A and B) on the microlens array MLAT,NS and its displacement hT,y regarding the microlens center position.

Transmission spot diameter on MLATHIT0 in %HIT0,hT in %PT0 in %PT0,hT in %rel. PTLoss(T0−T0,hT) in %
A: dT,MLAT=dμL56.52.186.540.253.5
B: dT,MLAT=2·dμL87.257.039.326.931.6

4.

Compensation of Spot Shadings

In Sec. 4, we show the drawback of partial shading of measurement spots, while scanning over the FoV. We describe how to avoid this drawback by an optimized optical design of a wedge prism array in the transmission path of the scanner system. Further, we evaluate the effect of spot shadings on the transmission power and consequently on the reachable measurement distance.

4.1.

Principle of Transmission Beam Cutoffs while Scanning over the Field of View

An inherent characteristic of scanner systems, including micro-optics for beam widening, is the emerging of transmission beam cutoffs at the aperture stop of the objective lens while scanning over the microlens array. The consequences are not only gaps between adjacent spots in the FoV. This also leads to a significant reduction of transmission power with coupled degradation of reachable measurement distance. Beam cutoffs at the aperture stop of the objective lens start at a system-specific incidence height of the transmission beam on the objective lens. This incidence height is transformed into a field angle, which we call the critical field angle wFoVcrit.. Based on these cutoffs, shadings or gaps hGap,Tr arise between adjacent measurement spots in the FoV and increase proportional to the tangent of the field angle wFoV [hGap,Trtan(wFoV)]. The consequence is possible nondetection of targets with higher spatial frequency and the reduction of measurement distance within higher field angles of the scanner system.

By realizing a maximum eye-safe transmission beam with a diameter equal to that of the objective lens LO, the cutoff already starts by the illumination of the first microlens aside the one to the optical axis. At the maximum field angle, the cutoff equates inherently always to half the spot diameter dTLO, which inevitably leads to a spot shading in the FoV of half the target diameter dTr. In Fig. 10(a), this is illustrated with the help of two exemplary transmission beams T0 and Tn.

Fig. 10

Influence of transmission beam cutoff on the residual spot height hSpot,Tr and the emerging gap height hGap,Tr between adjacent spots on the target. (a) Schematic functional principle of the beam-shaping section with and without MLAT/MWA combination. (b) Graph of the residual spot height in the FoV for an objective lens LO1 with a FoVh of ±23.5  deg without a MLAT/MWA combination at 100-m target distance. This is shown for two spot diameters dTLO=10.41  mm and dTLO=dLO=20.82  mm in the plane of the objective lens, where dLO is the diameter of the objective lens.

OE_58_8_087101_f010.png

Figure 10(b) shows the correlation between the increase of field angle wFoV and the reduction of the residual spot height hSpot,Tr up to 0.25- at 100-m target distance. On the basis of an objective lens with a FoVh=±23.5  deg and a diameter dLO=20.82  mm, we explain this characteristics for two spot diameters dTLO=10.41  mm and dTLO=dLO=20.82  mm in the plane of the objective lens. As we can see, in contrast to a fully illuminated objective lens, a spot diameter of 10.41 mm results in a critical field angle of 12.14 deg, which allows a gapless illuminated FoV of 56.26%. Hence, the percentage of the FoV, which is possible to be scanned without provoking shadings caused by these cutoffs, depends on the optical design of the beam-shaping section and the specified transmission spot diameter.

4.2.

Approach for Compensation of Spot-Shadings with Wedge Prisms

To avoid this substantial drawback, our optics solution includes a microwedge prism array MWA in addition to the microlens array MLAT. The microwedge prism array consists of one with the microlens array combined module or of two separate optical components optionally. In Figs. 10(a) and 11(a), the combined and the separate versions are illustrated. We design each wedge prism for one corresponding microlens exclusively. The wedge angle γw,n and the offset dw,n to the optical axis of the corresponding microlens depend on the radial distance hμL to the optical axis of the objective lens LO [see Fig. 11(c)]. The microwedge prism array refracts the divergent light cone of the transmission signal after the microlens for each scan position by exactly the amount that it can pass the objective lens LO without any cutoff at the aperture stop. We show in Fig. 11(b) the geometrical fundamentals for the calculation of the necessary tilt angle τtilt,n and the resulting wedge angle γw,n.

Fig. 11

Detailed functional principle of the microlens array/microwedge prism array MLAT/MWA combination. (a) Comparison of the optical transmission path between using a MLAT/MWA component and a single MLAT component without MWA. (b) Geometrical correlation of the necessary tilt angle τtilt,n and the resulting wedge angle γw,n of the wedge prism. (c) Schematic design of a MLAT/MWA component.

OE_58_8_087101_f011.png

From Fig. 11(c), we can calculate the radial distance of the aperture of each microlens hrμL,n to the optical axis of the objective lens LO with yμL,n=nμL,y·PitchMLAT,y, xμL,n=nμL,x·PitchMLAT,x, and hμL,n=xμL,n2+yμL,n2, which can be expressed as

Eq. (17)

hrμL,n=hμL,nrμL.
Here, nμL,x and nμL,y is the number of the particular microlens (in x- and y-directions), starting the numeration with the element on the optical axis of the objective lens LO.

The radial distance of the center of each particular microlens to this optical axis is given with hμL,n and the radius of the microlens with rμL=½·dμL. The necessary tilt angle τtilt,n of the marginal ray of each microlens is expressed as

Eq. (18)

τtilt,n=τμLτlim,n=tan1(rμLfμLrLOhrμL,naμL+fμL1+rμLfμL·rLOhrμL,naμL+fμL),
with τμL the maximum divergence angle of the light cone, assuming an entirely illuminated microlens. The tilt angle τlim,n describes the maximal angle, which can pass the objective lens LO with dLO=2·rLO. We can equate τtilt,n with the deflection angle of the wedge prism ϑw,n. With the assumption of γw,nαw,n=γw,n between the incidence angle αw,n of the transmission signal at the wedge prism and the wedge angle γw,n, we can write the equation as

Eq. (19)

γw,n=sin1((1+nw2)·sin(ϑw,n)±2·sin(ϑw,n)2C44·sin(ϑw,n)2+(1+nw2)2),
where C4=nw2sin(ϑw,n)2.

The refraction index of the wedge prism material is given with nw.

4.3.

Effects of Beam Cutoffs on Transmission Power and Residual Spot Height

We experimentally investigate the optimized micro-optical component, to demonstrate the improvement of a MLAT/MWA combination in a scanner system and to prove the theoretical calculations. The scanner setup provides a FoV of ±23.5  deg, a beam diameter of dTLO=10.41  mm, and a measurement distance of 0.90 m. Our tests reveal a high consistency between theoretical simulation and measurement [see Fig. 12(a)]. We can show that a setup with a microwedge prism array in the optical path increases the transmission signal in the FoV by the factor of 2.47 compared to that without such a component. In consequence, this theoretically leads to a rise of measurement distance (PTzTr)26 of about the factor of 1.57, assuming a fully illuminated target with a minimum diameter dTr.

Fig. 12

Measurement of transmission power PT over the field angle wFoV for an objective lens LO1 with a FoVh of ±23.5  deg. (a) PT over wFoV in comparison between simulation and measurement values. (b) Gradient of the loss of transmission power PTLoss over wFoV. The red dashed line gives the critical angle wFoVcrit., where the beam cutoff starts theoretically.

OE_58_8_087101_f012.png

The drop of the curve with MLAT/MWA combination is based on the incidence angle ϑw,n into the microlens, which increases the spot diameter with a cosine function in this plane. The result is less effective signal transmitting through the microlens. However, this can be avoided by rotating the microlens with the equal value to the value of ϑw,n or inverting the sequence of the components itself, which again leads to higher effort in appropriate lens design.

The gradient of the transmission power loss PTLoss in the FoV is given in Fig. 12(b). The curve shows the characteristic for an axial displacement in y-direction [similar to Fig. 10(a)] of a rectangular light cone to an objective lens with a round aperture stop. Because of the experimentally used rectangular microlens aperture, the critical field angle wFoVcrit.=8.67deg (red dashed line) is lower than the calculated angle of LO1 with a round microlens aperture, which is shown in Fig. 10(b). In the experiment, we see that the reduction of transmission power PT of the measured curve is higher compared to the simulated curve. The reason for this is the telecentric characteristics of the objective lens LO utilized in the experiments. In such a lens, the clear aperture decreases significantly with higher scan angles wFoV. This also causes the increasing deviation between simulation and measurement curves with increasing wFoV. We can avoid this issue with appropriate lens design of the objective lens LO as well.

In Fig. 13, we try to clarify the effects of transmission beam cutoff at the aperture stop of the objective lens LO on the residual spot height hSpot,Tr, which is projected on the target at 0.45-m target distance. In the first row, we see the decrease of hSpot,Tr for defined field angles wFoV=0.00  deg, 9.09 deg, 16.51 deg, and 22.16 deg without MLAT/MWA combination. Starting at the critical field angle wFoVcrit, the value of hSpot,Tr shows the behavior of hSpot,Trtan(wFoV). In the second row with the MLAT/MWA combination, the upper cutoff of the corners arises from rounding errors between the calculated radial distance to the optical axis and the utilized wedge prism with given wedge angle γw,n. The lower cutoff starting at 16.51 deg and increasing significantly at 22.16 deg is based on the already-mentioned telecentric characteristics of the utilized objective lens LO.

Fig. 13

Effect of transmission beam cutoff on the spot geometry and the residual spot height hSpot,Tr on the target. Here, the figures show the comparison between the spots for several field angles wFoV at 0.45-m target distance with and without a microlens array/microwedge prism array combination (MLAT/MWA).

OE_58_8_087101_f013.png

In Fig. 14, we show the measurement of the decreasing residual spot height, while scanning up to a field angle of 22.16 deg without a microwedge prism component. Here, the measurement result shows good consistency to simulated values. The reduction of the spot height starts at a smaller field angle with 10.73 deg compared to that of the simulation with 12.14 deg. This is based on a minimal smaller diameter of the aperture stop of the objective lens used in the experiments as assumed in the simulation.

Fig. 14

Measurement of spot height hSpot,Tr over the field angle wFoV for an objective lens LO1 with a FoVh of ±23.5  deg without a MLAT/MWA combination. The red dashed line gives the critical angle wFoVcrit., where the beam cutoff starts theoretically.

OE_58_8_087101_f014.png

5.

Conclusion

For an eye-safe scanner system, we show principal limitations regarding the angular resolution and the maximum possible diameter of the transmission beam. The upper limit for the transmission beam diameter is about 20 mm with a diameter of microlenses of 99.6  μm. This is calculated for a micro-optics-based scanner system with an angular resolution of 0.286 deg. Such an angular resolution equates to the resolution of a target with the diameter 0.5-m at 100-m target distance, assuming an f-number of the microlens of 1. The system allows MPEs of 4  kW peak power within 2.5 ns FWHM at 900-nm wavelength. Here, a FoV of ±28.5  deg is reachable with a diameter of the objective lens of 20.82 mm.

To suppress spot position noise in the FoV, we present an approach of an optimized microlens array with carefully designed aperture stops in the transmission path. We are able to suppress a mirror-deflection noise with an angular uncertainty of 0.013 deg. This equates to a suppression of a spot position error of about 0.25-m at 100-m target distance. With our approach, we additionally improve the transmission power loss by about 21.6% and the transmission spot homogeneity by about 54.9%, depending on the level of spot position error on the microlens array. The power consumption of a setup with position-noise-suppressing microlens design is about a factor 2.19 higher compared to a setup without this design.

The scanner system, presented in this paper, allows a small rotating mirror and provides a high spot position accuracy of measurement spots in the FoV without any shadings in between. With the combination of a microlens array and a microwedge prism array in the optical transmission path, we can show a nearly perfect compensation of transmission beam cutoffs at the aperture stop of the objective lens. In consequence, we avoid the resulting shadings between adjacent spots in the FoV up to half the specified target diameter, which is to be resolved. This realizes not only a gapless illuminated FoV but also a significant increase of transmission signal by a factor of 2.47. Theoretically, this leads to a rise of measurement distance of about a factor of 1.57, derived from the radiometric inverse square law.26

Our experimental setup offers an objective lens with 20.82 mm diameter and a FoV of ±23.5  deg at 0.90-m target distance. The FoV can be scanned with a deflection angle of the oscillating mirror of about ±3.95  deg at 75-mm focal length of the lens LT. We utilize a microlens with an f-number of 2.4, which generates a spot diameter of 10.41 mm at the system exit plane on the objective lens. The MPE of this setup is about 1 kW transmission peak power.

Acknowledgments

The presented work has been developed in collaboration with Robert Bosch GmbH and the Clausthal University of Technology. The authors sincerely acknowledge the support and want to thank all contributors and colleagues for their helpful inputs and discussions to realize this project.

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Biography

Jürgen Hasselbach received his diploma degree in mechatronics and microsystems from Heilbronn University of Applied Sciences, Germany, in 2010. From 2010 to 2016, he worked at the Robert Bosch GmbH as development engineer in optical testing for video-based automotive driver assistance systems. Since 2016, he has been a PhD student at Clausthal University of Technology in cooperation with the Robert Bosch GmbH. The subject of his work is the development and investigation of a LiDAR scanner module with large sending/receiving apertures and high angular position accuracy.

Siegwart Bogatscher received his diploma in electrical engineering from the University of Karlsruhe (TH), Germany, in 2009. He was a PhD student at the Karlsruhe Institute of Technology (KIT) from 2009 to 2014 and received his doctorate degree in electrical engineering. In 2015, he joined Robert Bosch GmbH. As a system architect, he works in the field of LiDAR development for autonomous driving applications.

Christian Rembe received his diploma in physics from the University of Hanover, Germany, in 1994. From 1994 to 1999, he was a PhD student at the University of Ulm, Germany, where he earned a doctorate degree in electrical engineering. In 1999, he joined the Berkeley Sensor & Actuator Center at the University of California, Berkeley, as a postdoctoral research engineer with the support of a Feodor-Lynen-Scholarship of the Alexander von Humboldt-Foundation and a UC Berkeley MacKay-Lecturer Fellowship. From 2001 to 2014, he was the manager of the Development Optics group at the Polytec GmbH in Waldbronn, Germany. Since 2015, he has been a professor for applied metrology at the Clausthal University of Technology, Germany. His interests are optics, optical measurement techniques, sensors, and microelectromechanical systems.

© The Authors. Published by SPIE under a Creative Commons Attribution 4.0 Unported License. Distribution or reproduction of this work in whole or in part requires full attribution of the original publication, including its DOI.
Juergen Hasselbach, Siegwart Bogatscher, and Christian Rembe "Laser scanner module with large sending aperture and inherent high angular position accuracy for three-dimensional light detecting and ranging," Optical Engineering 58(8), 087101 (7 August 2019). https://doi.org/10.1117/1.OE.58.8.087101
Received: 11 March 2019; Accepted: 2 July 2019; Published: 7 August 2019
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