dual field-of-view midwave infrared optical design and

Dual field-of-view midwave infrared optical designand athermalization analysis

Chih-Wei Kuo,1,* Chih-Lung Lin,2 and Chien-Yuan Han3

1Electro-Optics Section, Materials and Electro-Optics Research Division, Chung-Shan Institute of Scienceand Technology, Lung-Tan 325, Taiwan, China

2Department of Electronic Engineering, Hwa Hsia Institute of Technology, Taipei 235, Taiwan, China3Department of Electro-Optical Engineering, National United University, Miao-Li 360, Taiwan, China

*Corresponding author: [email protected]

Received 2 April 2010; revised 31 May 2010; accepted 2 June 2010;posted 2 June 2010 (Doc. ID 126354); published 23 June 2010

A step-zoom and reimaging structure were utilized to construct a dual field-of-view optical design forhigh-magnification switching in the 3–5 μm spectral band. The design has a flexible optomechanical lay-out, which means it can be utilized for multipurpose applications. The effects of the surrounding envir-onmental temperature and axial gradient temperature are analyzed using the concept of thermalresistance, and the thermal compensation is discussed. A description of the zooming mechanism andoptomechanical control is offered. © 2010 Optical Society of AmericaOCIS codes: 110.3080, 220.3620.

1. Introduction

The applications of infrared dual field-of-view(DFOV) optical systems have been widely discussed[1–5] for both themidwave infrared (MWIR) spectrum(3–5 μm) and longwave infrared spectrum (8–12 μm).These optical systems utilize the wide-angle field ofview (WFOV) to search the scenery and the narrowangle (NFOV) mode to identify the target of interestclose up. The algorithm for assessing a surveillanceimage is derived from the Johnson criteria [6] andthen used to deduce the target being detected, recog-nized, or identified [7]. Variations in the characteris-tic size of a particular object, and the distance fromthe sensor to the target plane, both result in the dif-ferent magnification ratio requirements obtained bythe computation of the Johnson criteria. Previously,theDelano diagramhas been used to aid in the designof zoom lenses [8], composed using the height of theparaxial marginal and chief rays as the longitudinaland transverse coordinates, respectively. The para-

xial analysis of mechanically compensated zoomlenses can be expressed in terms ofGaussian brackets[9,10]. The focus of the zoom lens can be solved bydescribing the relationships among focal length, lensposition, ray height, and direction in relation to thematrix. The zoom process can be expressed by aunified varifocal differential equation with a stableimage plane being the constraint condition [11]. Ther-mal imaging systems for target-tracking purposes,such as rocket and missile launching applications,require WFOV systems for monitoring and NFOVsystems for identification. DFOV infrared optical sys-tems should have FOV switching without target loss.They can be classified as a subset of continuous-zoomstructures that utilize only axial steps to move lensesbetween two extreme magnifications. A nonaxialmoving DFOV variant, called the rotating-in scheme,was developed by combining reverse-telephoto andtelephoto structures with/without two separate lensgroups controlled by a rotation mechanism. This isdistinct from the step-zoom scheme [12].

Most infrared semiconductor detectors are cryo-genically cooled and assembled in a thermally insu-lated Dewar flask. This is necessary to achieve the

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1 July 2010 / Vol. 49, No. 19 / APPLIED OPTICS 3691

maximum signal-to-noise ratio and avoid anomalousimages. A special baffle, called a cold stop, located in-side the Dewar flask, causes the light cone to strikethe focal plane array (FPA), and the FPA records theobjective space energy exclusively. It has been deter-mined that 100% cold-stop efficiency is required [13].This can be accomplished by the aperture stop andcold stop coinciding. The optical system is still opera-tional if the effective focal length (EFL) is small.However, to identify or classify a distant but smalltarget, a longer EFL is essential. The layout of theaperture stop inside the Dewar flask will make thediameter of the front end of the lens larger thanthe entrance pupil, because of the off-axis light coneaccommodation. The only way to shrink the aperturesize of the front end of the lens while maintaining100% cold-stop efficiency and offering a longer EFLis through reimaging. The extra relay lens creates animage of the aperture stop located on the front endlens that is equal in diameter to the entrance pupil.The reimage element is the primary objective, fol-lowed by the intermediate focal plane, relay lens,cold/aperture stop, and FPA.

Stray light caused by double reflection in the visi-ble lens can create ghost images, especially in a scenewith a signal difference source against a dark back-ground. A high transmission rate of antireflectioncoating in the IR spectrum is harder than in the visi-ble one; hence, an IR system is more vulnerable todouble-reflection stray light. Besides, the cryogeni-cally cooled detector acts as a strong light in a darkscene because of its low temperature compared withthe warmer surroundings of the lens barrel or the ob-served target. Therefore, an infrared FPAmay reflecta single reflection of its own image. This is called thenarcissus effect [14,15], and this retroreflected non-uniformity thermal image reduces the contrast of adim object. In a scanning system, the narcissus effectis noticed as a retroreflected ghost image that moveson the focal plane as the scanning mirror rotates. Thestarring array system can negate the narcissus effectif the optical system meets the requirements of non-uniformity calibration, a constant surrounding tem-perature, and no lens movement. Otherwise, thenarcissus effect is monitored by an optical designprocess.

Temperature change in the operating environmentcan result in variation of the physical properties of anoptical assembly (i.e., lens thickness, air space, refrac-tive index, structural dimensions) [16]. These effectscan cause the system to lose focus or elements to be-come misaligned. Temperature gradients may causetilting in the axial or radial directions, reversing theeffects of homogenization on the materials and dis-torting the optical surface. The quality of the final im-age can become degraded as the adverse temperatureeffects increase. Most materials for infrared refrac-tive lenses have high rates of index variation withtemperature, which can result in rapid focus shifts.Degradation from thermal variations in an infraredsystem is an order of magnitude higher than in a visi-

ble system. The methodology and process for stabiliz-ing the optical qualities by temperature compensa-tion is called athermalization.Active athermalizationis appropriate for quick responses to overcome rapiderror inputs while operating in a thermally dynamicenvironment [17]. To facilitate athermalization, amo-torized lens adjusts the position and thermal sensorsmeasure the temperature difference. The lens andthe sensor are controlled by a programmed microprocessor.

An advantageous feature of the DFOV infrared op-tical system is the incorporation of motorized me-chanical parts that allow it to travel quickly tocompensate for the variation of the optical group overan exact distance. In other words, the optomechani-cal design can ensure the functioning of the opticaldesign. The mechanical layout of the zoom system isadapted to meet overall system requirements. Inno-vations and inventions have led to the developmentof numerous patents. For example, two coaxial lenssleeves have been assembled together. With this typeof device, the translation of optical elements is en-abled by the relative rotation of the sleeves. Thecam grooves provide a predetermined locus [18]. Thebarrel is made more compact and portable when notin operation by the movement of a thrust memberback and forth in the optical-axis direction [19].The diameter of the whole barrel can be reducedby replacing the coaxial sleeve layout with an off-axiscam attached to the lens rim [20]. The traditionalcam groove located on the cylindrical surface is re-placed by the plane surface [21]. The cam providesa reliable lens walkway but the mechanical manufac-turing process is expansive and complex. Thus, thelens movement path can be controlled by a pro-grammed motor utilizing a spur gear to drive thefeed bar [22]. The optomechanical alignment mustensure that the line of sight (LOS) remains unaf-fected during the change in the FOV, especially forautomatic target-tracking applications. The methodsfor LOS stabilization commonly used to counteractjitter introduced from the optical system platformand surroundings can be categorized in three types:software, platform, and steering [23].

2. Lens System Scheme

A SELEX HgCdTe FPA sensor (F=4, 384 × 488 pixelsand 20 μm square pixel size) was used in this study.The optical design offers a dual EFL of 20=250mm.This specification can improve the broad scene whilestill giving high enough resolution for most infraredoptical system requirements. This is a reimaging typesystem that allows for reduction of the lens apertureand for compact instrument space. To ensure imagequality, we must strictly consider a better minimumresolvable temperature difference (MRTD). The poly-chromaticmodulation transfer function (MTF)withinthe 3–5 μm spectrum approaches the diffraction limitbeneath the Nyquist spatial frequency.

In a Gaussian design, a primary objective is imple-mented to generate the system optical power and to

3692 APPLIED OPTICS / Vol. 49, No. 19 / 1 July 2010

focus on the intermediate focal plane under infiniteconjugate conditions. This is reimaged by a relay lensto the final focal plane (i.e., the cryogenic detector ar-ray). This layout must ensure that the system aper-ture stop coincides with the Dewar’s cold stop for100% cold-shading efficiency. The NFOV mode is re-quired for the layout of the telephoto objective lenssystem. The unit power relay lens helpsmake for eco-nomical saving of the axial length. Adjustment ofboth the focal length of the telephoto objective sys-tem and the relay lens allows the entrance pupilto be located on the primary objective, which canminimize the primary lens diameter. Wandering ofthe pupil, caused by off-axis beam accommodation,can be eliminated [12,13]. These calculations werecarried out with a Gaussian imaging equation andparaxial yu ray tracing:

n0ju

0j ¼ njuj − yjϕj; ð1Þ

yjþ1 ¼ yj þ tju0j; ð2Þ

where u is the tangent value of the angle between theoptical axis and the ray pencil; y is the contact heightof the ray; t is the thickness of the air spacing; and ϕis the optical power of the lens. The subscript symbolrepresents the optical surface number, and the super-script prime represents the ray after refraction. Onecomponent of the lenses composing the telephoto ob-jective system is able to move to change the magni-fication and the other component to eliminate thefocal shift. Consequently, we were able to constructa WFOV with a reverse-telephoto layout. This DFOVoptical system was comprised of four sequentialgroups: the focusing group, variation group, compen-sation group, and relay group. The lens power of eachgroup was positive, negative, positive, and positive,respectively [24]. The Gaussian arrangement isshown in Fig. 1. The EFL of the optical system canbe determined by the following equation:

f system ¼ ðϕ1 þ ϕ2 þ ϕ3 − t1ϕ1ϕ2 − t1ϕ1ϕ3 − t2ϕ1ϕ3

− t2ϕ2ϕ3 þ t1t2ϕ1ϕ2ϕ3Þ−1 ×mrelay; ð3Þ

where f system is the EFL of the optical system, andmrelay is relay lens magnification. Theoretically,switching the EFL can be accomplished by move-ment of the varying and compensating groups.

3. Optimization of Image Quality

Many Gaussian solutions, including optical powerandair space, canbe foundbefore aberrationsand fea-sibility are taken into consideration. The reasonableinitial arrangement of optical power and air space foreach group and the yu ray-tracing data are listed inTable 1. In NFOV mode, the first lens diameter isequal to the entrance pupil, and the wandering ofthe pupil is not existent. The material propertiesand shape factors of the lens are set accordingly. Sili-con is used primarily for its high refractive index,which is advantageouswith respect to aberration con-trol. Germanium and silicon are applied to create anachromatic air-spaced doublet, because their Abbenumbers have great difference. The total monochro-matic Seidel aberrations and lateral/axial chromaticaberrations in lenseswere evaluated. This evaluation

Table 1. NFOV and WFOV Paraxial Data

NFOV Paraxial DataSurf Type Power (mm−1) Thickness (mm) Y Marginal (mm) U Marginal

OBJ Standard Infinity 0.000 0.0001 Paraxial 0.004 170.0 31.250 −0:1252 Paraxial −0:025 80.0 10.000 0.1253 Paraxial 0.0125 160.0 20.000 −0:1254 Standard 80.0 0.000 −0:1255 Paraxial 0.025 80.0 −10:000 0.125

IMA Standard 0.000 0.125

WFOV Paraxial DataSurf Type Power (mm−1) Thickness (mm) Y Marginal (mm) U Marginal

OBJ Standard Infinity 0.000 0.000STO Paraxial 0.004 38.57 2.500 −0:0102 Paraxial −0:025 264.00 2.114 0.0433 Paraxial 0.0125 107.43 13.429 −0:1254 Standard 80.00 0.000 −0:1255 Paraxial 0.025 80.00 −10:000 0.125

IMA Standard 0.000 0.125

Fig. 1. (Color online) Gaussian arrangement.


wasconductedwith theZEMAXcommercial software,after the variables, zoom positions, and system con-straints were assigned. The weighting functions ofthe FOV were set equally at the on axis, 0.5, 0.707,0.866, and at the full FOV. The MWIR spectra (3, 4,and 5 μm), were distributed as the same weightingfunctions. The location of the Dewar cold stop was de-fined by the system aperture. The layouts of the opti-cal system and aperture/cold stop position are shownin Fig. 2. All lenses are housed in one barrel exclu-sively, and the other elements (window, filter, stop,and FPA) are assembled as a radiometry detectorunit. The requirements of optical manufacturing andmechanical assembly were translated into the pro-gram’s merit function editor. The ZEMAX softwareutilized the damped least-square algorithm to mini-mize the merit function [25]:

MF ¼Xj

WjðVj − TjÞ2; ð4Þ

where W is the weighting function; V is the currentvalue; andT is the target value. The subscript symbolrepresents the operand item number. As a startingpoint for optimization, the rms spot radius is selectedas a merit function. During the final stage of optimi-zation, wavefront errorwas selected as systemperfor-mance approached the diffraction limits. The ray fansare shown in Figs. 3(a) and 3(b).

Backward ray tracing from the cryogenically cooleddetector can determine the surface of the contributingretroreflection ghost. However, this procedure is time

consuming and labor intensive. It is more efficientand of greater benefit to simultaneously design an op-tical lens to control the narcissus effect. The ratio ofthe detector area (A) multiplied by the average emis-sivity (ε) to the area of the ghost defines the narcissusintensity of a single surface, and this ratio is alsonamed as cold return (CR ¼ εA=ðπ½4yniðF=#Þ�2Þ, de-tail equation and figure referring [14]). Summingall surface ratios, CR represents a system’s narcissus.The product of three geometric optical parameters,yni, is proportional to the ghost area. Here, F/# isthe numerical aperture, n is the refractive index,and i is the reflective marginal ray angle of the inci-dence surface. If the value of yniwas larger, the inten-sity of the narcissus was lower.

A narcissus intensity for the whole detector areaand ghost variations across every pixel can be effi-ciently controlled by constraining the absolute valuesof yni (Table 2) using an optical design program.How-ever, these restrictions confined the development ofthe lens shape factor. This could be a problem forthe required diffraction-limited performances. There-fore, a trade-off is needed. In addition, a lens with ahigh-efficiency antireflection thin film coating (the

Fig. 2. (Color online) (a) NFOV, EFL ¼ 250mm. (b) WFOV,EFL ¼ 20mm. (c) Position of aperture/cold stop.

Fig. 3. (Color online) (a) NFOV, ray fan plot. (b) WFOV, ray fanplot.


ghost strength depending on the residual reflection atthe offending lens surface) and sharp edge spectrumbandpass filter inside the Dewar flask are helpful torelax this anomalous image.

4. Athermalization Analysis

Thematerial properties and the characteristic config-urations of optical elements andmechanical parts canchange with variations in the surrounding tempera-ture. A methodology must be developed to compen-sate for the optical system performance and torecover the degradation of the image quality. Mostlenses are made of a brittle material, so the surfaceradius departs from the original design value. How-ever, the major problem caused by temperature var-iation is the term for differentiating the refractiveindex with temperature (dn=dT). This is particularlytrue in an infrared system but less so in a visible sys-tem. Reference [17] lists the dn=dT of most visible op-tical materials as falling between 3 × 10−6=°C and10 × 10−6=°C (such as BK7, dn=dT ¼ 3:6 × 10−6=°C),and that of infrared materials as between 30 ×10−6=°C and 400 × 10−6=°C (such as germanium,dn=dT ¼ 396 × 10−6=°C). Generally, the difference ofdn=dT in infrared materials is an order of magnitudehigher than in visiblematerials. Therefore, it is neces-sary to consider thermal compensation in infraredspectrum optical designs.

There are two distinct thermal compensationmethodologies: passive and active. In most visiblesystems, as the temperature rises, the elongation ofthe barrel length is greater than the growth of theEFL. The simplest way to compensate for this is tochoose a mechanical material that allows the ther-mal expansion of the barrel to be close to the opticalthermal defocusing. This is considered passive com-pensation. Another passive technique is the bi-metal

technique, used in both serial and reentrant layoutsso that the thermal variation of the barrel matchesthe optics [26]. In infrared spectral imaging applica-tions, thermal defocusing is almost larger than thebarrel expansion allows. Therefore, appropriate me-tal selection is vital. However, passive compensationis only suitable for steady temperature states. In athermally dynamic environment, a motorized com-pensation lens can be used to slightly displace theoriginal position. Compensation is calculated witha specific algorithm, which can satisfy both steadyand transient thermal states. This is called activecompensation.

Most infrared optical system applications requirehigher target resolutions; therefore, a longer EFL isessential. The total length of the infrared optical as-sembly is obviously longer than that of most commer-cial visible systems, even thoseadopting the telephotostructure. The temperature difference between theoptical system’s front and bottom sides cannot be ne-glected in an infrared lens integrated into a gimbalspod or vehicle turret, where only the first lens surfaceis exposed to the outside environment. The axial tem-perature gradientmust be taken into consideration incases where the optical system suffers from variationin environmental temperature. Thermal effects notonly arise from shifts in the soaking temperature,but also from the existence of the axial temperaturegradient. One-dimensional thermal resistance analy-sis can be utilized [27] to evaluate axial temperaturedistribution:

Rj ¼lj

Ajkj; ð5Þ

H ¼ ΔTΣRj

; ð6Þ

where R is the thermal resistance, l is the character-istic length, A is the heat flux crossing area, k is thethermal conductivity, ΔT is the temperature differ-ence, andH is the heat flux. The subscript symbol re-presents the item number of the thermal resistor. Inthis study, we employed a silicon (monocrystalline) al-loy for the lens and a germanium (monocrystalline)and aluminum alloy (No. 6061) for the barrel. Thethermal conductivities of these materials are shownin Table 3 (polycrystalline and the other aluminumalloy numbers will be different tabulated values).The thermal resistance diagram is shown in Fig. 4.The axial heat flux is primarily conducted by the bar-rel parts because the thermal resistance of air is sev-eral orders higher than that of aluminum alloys,germanium, or silicon. The axial thermal expansion

Table 2. Absolute YNI Value

Surf YNI (NFOV) YNI (WFOV)

1 4.55 0.032 4.92 0.033 4.97 0.034 1.02 0.015 0.58 0.096 0.15 0.087 0.02 1.218 0.11 1.499 0.09 1.5810 0.05 0.4011 0.05 0.3012 0.23 0.2313 0.94 0.9414 0.96 0.9715 0.31 0.3116 0.80 0.7917 0.05 0.0518 0.33 0.3319 0.32 0.3220 0.30 0.3021 0.30 0.29

Table 3. Material’s Thermal Conductivity

Aluminum Germanium Silicon Air

Thermal Conductivity(W/mK)

180 59 163 0.026


of the barrel can be determined by calculating the ax-ial temperature distribution with the following equa-tion:

TðzÞ ¼ Tmin þ�Tmax − Tmin

L

�z; ð7Þ

whereTðzÞ is the axial temperature distribution,TminandTmax are the temperature gradients on both sides,L is the length of the barrel, and z is the coordinate forthe axial direction. Moreover, each lens rim ismounted between the cell and the retainer. The airsurrounding the lens provides better thermal insula-tion. Therefore, the lens radial temperature distribu-tion is assumed to be uniform and equal to the cell’stemperature. For the thermal setup, each lens surfaceradius, thickness, and index at a specific temperaturecan be calculated using ZEMAX’s multiconfiguration

function. Finally, these data are input to the opticalsoftware for ray tracing for the whole optical system.The performance of the optical system given the axialtemperature gradient can be analyzed.

In the present study, the reference temperature isalways 20 °C, and the other soaking temperature va-lues are comparedwith this one. In the beginning, theMTF curves for the soaking temperature (0 °C, 20 °C,40 °C, and 60 °C) are plotted in Figs. 5(a) and 5(b). Theoff-axis curves smoothly degrade as temperature de-parts from 20 °C, for the thermal variations of opticaland mechanical materials are getting larger as thetemperature changes. The on-axis MTF is discussed,because the optical axis is boresighted for aiming.There is little thermal effect to the on-axis MTFcurves [Figs. 5(c) and 5(d)] even at higher frequency.This could be the paraxial results being affectedslightly by the soaking temperature. The on-axiscurve of 40 °C seems a little bit higher than the others,forweighting functions of all FOVare the sameduringoptimization. However, the tangential and sagittalcurves spread out as the temperature ranges fartheraway from 20 °C. The distortion and field curvature

Fig. 4. Thermal resistance diagram.

Fig. 5. (Color online) (a) NFOVandMTF results at 0 °C, 20 °C, 40 °C, and 60 °C soaking temperature. (b) WFOVandMTF results at 0 °C,20 °C, 40 °C, and 60 °C soaking temperature. (c) NFOV and MTF on-axis results for 0 °C, 20 °C, 40 °C, and 60 °C soaking temperature. (d)WFOV and MTF on-axis results 0 °C, 20 °C, 40 °C and 60 °C soaking temperature.


depending on FOV show insignificant changes at dif-ferent soak temperatures.

The temperature gradient is defined by the tem-perature at the front side of the barrel minus the tem-perature at the bottom of the barrel. The lens barrelbottom side temperature is always 20 °C [as Fig. 2(c)indicates, the window, filter, stop, and cryogenic FPAare evacuative as a stand-along unit, which is sepa-rated from the lens barrel by air space.). The MTFcurves for axial gradient temperatures of −20 °C,þ20 °C, and þ40 °C are shown in Figs. 6. Thermaldegradation was manifest in both the NFOV modeand the WFOV mode, which was especially sensitiveto the existence of the axial temperature gradient.The variations of distortion and field curvature de-pending on FOV are negligible as gradient tempera-ture alterations.

As discussed above, this optical system is consti-tuted of four groups, one of which is the compensat-ing group, counterbalancing any defocus originatingfrom variation group. Consequentially, the neutrali-zation of thermal defocusing can be analyzed. Thedisplacement of the compensating group can counter-

act the MTF degradation introduced by the axialtemperature distribution, as shown in Figs. 7. Thesecurves exhibit the athermalized benefit of the com-pensating group. Although efficient for the WFOVmode, the NFOV mode can still tolerate some resi-dual unsolved thermal defocusing. Because of thechromatic aberration in the NFOV mode, the ther-mal defocus cannot be eliminated completely by themotion of the compensation group. The curves of on-axis polychromatic rms spot radius plotted againstgradient temperature are shown in Fig. 8. It can beseen that the polychromatic spot radii in the NFOVmode were larger than those in the WFOV mode. Itindicates that the NFOV mode suffered temperaturegradient effect more than the WFOV mode. Duringthis EFL 20=250 DFOVapplication, the WFOV helpsto search one object (assuming its image spatial fre-quency f image), and the NFOV identifies the same ob-ject (image spatial frequency being f image divided bythe zoom ratio, 12.5). This optical design can offer abetter modulation at f image=12:5 in NFOV than thatat f image in WFOV (cf. Fig. 7), even suffering the ad-verse temperature changing effects. The chromatic

Fig. 6. (Color online) (a) NFOV and MTF results at −20 °C, þ20 °C, and þ40 °C temperature gradient. (b) WFOV and MTF results at−20 °C, þ20 °C, andþ40 °C temperature gradient. (c) NFOVandMTFon-axis results at −20 °C, þ20 °C, andþ40 °C temperature gradient.


aberration coming from thermally dependent indexvariation is residual somewhat. However, there area limited number of lens materials that are usedin MWIR spectral bands, and the common solutionis a silicon–germanium air-spaced achromatic doub-let. The image quality has been satisfied in most ap-plications.

5. Optomechanical Control and Layout

The optomechanical actuation of the varying andcompensating groups was planned utilizing two inde-pendent feed bars driven by programmed DC motors[28]. Amicroprocessor, encoder, andmotor,making upa closed-loop feedback control system,were utilized topropel eachmoving lens group to the desired position.The proportional-integral-derivative (PID) algorithm(which is part of a fully developed andmature technol-ogy) acts as the controller’s mathematical back-ground. An RS232 interface and proper protocolsplay the communication role. The results show thatthe field of view can be switched as required.

Thermal sensors were attached to the front andbottom sides of the barrel [29], so that the axial

temperature gradient could be interpreted by the mi-croprocessor. To counteract thermal defocus, the dis-placement of the compensation group was carried out

Fig. 7. (Color online) (a) NFOVand compensated MTF results at −20 °C, þ20 °C, and þ40 °C temperature gradient. (b) WFOVand com-pensated MTF results at −20 °C, þ20 °C, and þ40 °C temperature gradient. (c) NFOV and compensated on-axis MTF results at −20 °C,þ20 °C, and þ40 °C temperature gradient. (d) WFOV and compensated on-axis MTF results at −20 °C, þ20 °C, and þ40 °C temperaturegradient.

Fig. 8. (Color online) Temperature gradient, NFOV, WFOV, andcompensated spot size.


according to the temperature distribution. And thisprocedure was arranged by the automatic controlprogram stored in the memory of the microprocessor.Consequently, stand-alone athermalization wasachieved.

The present design is intended for multipurposeapplications. The optomechanical layout [30–32]can be straight, L, or U-shaped [Figs. 9(a) and 9(b)],given the electro-optical (EO) system’s mounting re-quirements for portable finders, vehicle turrets, orgimbal pods. The folding mirror offers optical pathturning ability but requires careful placement toavoid reimaging the mirror surface’s defect to theFPA if it were placed right on the intermediate focalplane. The folding mirror provides an extra benefitfor aligning the optical axis and barrel axis duringassembly. It can also be steered to act as a line-of-sight stabilizer to counteract jitter arising from theEO system’s platform or surrounding environment.A precision gyroscope, quick response actuator, andcomplex digital signal processing algorithm areneeded to facilitate state-of-the-art performance.

6. Conclusion

A DFOV infrared optical design is described, andconstructed of positive, negative, positive, and posi-tive optical power groups. An intermediate focal

plane is located between the first three groups andthe final group. This type of reimaging can ensurethat the entrance pupil matches a narrow field ofview to produce a cold stop. The three previousgroups execute the field-of-view switching function.Their EFL times the final group’s magnification de-termines the EFL of the whole optical system:20=250mm. After optimization, the MTFapproachesthe diffraction limit.

Different soaking temperature and axial gradienttemperature situations are analyzed. The concept ofthermal resistance is employed to simplify the com-plex calculation. Uniform temperature changes donot lead to deterioration of optical performance whenthe elongation of the barrel and variation in thermalfocus are similar. However, the axial temperaturegradient results in a decline of quality of this opticalsystem. To compensate for defocusing while zooming,thermal defocus is compensated for by a shift in thethird optical group.

To activate motion of the desired optical group forEFL switching, the PID algorithm drives the motor-ized feed bar and the encoder feedback composes aclosed loop control system. The microprocessor moni-tors axial temperature distribution based on thethermal sensor for execution of the compensatingmovement for athermalization. To ensure that thisoptical design can be applied for multipurposes, theoptical path can be folded with an extra planar mir-ror. And the folding mirror can also benefit lensalignment and LOS stabilization.

The author acknowledges the contributions of col-leagues in the EO Section, Materials and Electro-Optics Research Division, Chung-Shan Institute ofScience and Technology.

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