INST, OF STAND & TECH R.I.C NIST

PUBLICATIONS

AlllDS DHTSTM

nited Stc

Technolog>

National Insi

NIST Special Publication 905

Techni

Symposium

MeasureI

Sponsored by the National Insi

in cooperation with the IEEE

and the Optical

NIST Special Publication 905

Technical Digest—Symposium on Optical

Fiber Measurements, 1996

Digest of a symposium sponsored by the

National Institute of Standards and Technology

in cooperation with the IEEE Lasers and Electro-Optics Society

and the Optical Society of America

Edited by

G.W. Day

D.L. Franzen

P. A. Williams

Optoelectronics Division

Electronics and Electrical Engineering Laboratory

National Institute of Standards and Technology

Boulder, Colorado 80303-3328

October 1996

U.S. DEPARTMENT OF COMMERCE, Michael Kantor, Secretary

TECHNOLOGY ADMINISTRATION, Mary L. Good, Under Secretary for Technology

NATIONAL INSTITUTE OF STANDARDS AND TECHNOLOGY, Arati Prabhakar, Director
National Institute of Standards

and Technology

Special Publication 905

Natl. Inst. Stand. Technol.

Spec. Publ. 905

220 pages (October 1996)

CODEN: NSPUE2

U.S. Government Printing Office

Washington: 1996

For sale by the Superintendent of

Documents

U.S. Government Printing Office

Washington, DC 20402-9325
PREFACE

Measurements of polarization mode dispersion (PMD) and nonlinear processes in optical fibers are two of the major topics in this digest of papers presented at the ninth Symposium on Optical

Fiber Measurements, held October 1-3, 1996, at the laboratories of the National Institute of

Standards and Technology in Boulder, CO. Summaries of all the papers presented at the

Symposium-10 invited and 39 contributed-are included.

The statistical nature of PMD complicates its measurement and makes verification of accuracy

difficult. This is reflected in the content of the papers, many of which focus on comparisons of

measurement techniques. In wavelength division multiplexed systems employing optical

amplification, high powers can lead to problems from nonlinear processes such as Brillouin

scattering and four-wave mixing. The importance of these issues is seen by the inclusion of

several papers regarding the measurement of nonlinear coefficients and effective area.

The Symposium continues to be a prime gathering place for specialists in the characterization of

optical fiber and related components. Strong international participation has been maintained, as about two thirds of the papers in this digest originated outside of the United States.

G.W. Day

D.L. Franzen

P.A. Williams

Boulder, Colorado

October 1996

Except where attributed to NIST authors, the content of individual sections of this volume has not been reviewed or edited by the National Institute of Standards and Technology. NIST therefore accepts no responsibility for comments or recommendations

therein. The mention of trade names in this volume is in no sense an endorsement or recommendation by the National Institute of Standards and Technology.

iii
SYMPOSIUM COMMITTEE

D.L. Franzen, NIST, General Chair

G.W. Day, NIST, Program Chair

M. ArtigUa, CSELT

A. Barlow, EG&G

S. Fleming, University of Sydney

T.A. Hanson, Coming

B.L. Heffner, Hewlett-Packard

R.K. Hickemell, NIST

R.B. Kummer, Lucent Technologies

P.S. Lovely, Photon Kinetics

S. Pollitt, NPL

W.A. Reed, Lucent Technologies

P.R. Reitz, AMP

C. Saravanos, Siecor

P.A. Williams, NIST

H. Yajima, Electrotechnical Laboratory

H.H. Yuce, Bellcore

iv
CONTENTS

PREFACE iii SYMPOSIUM COMMITTEE iv

Characterization of optical components for D-WDM applications (invited)

J. Dupre, Hewlett-Packard Company 1

Design and measurement of optical fibre gratings (invited)

Raman Kashyap, BT Laboratories 7

Chirp determination of fiber Bragg gratings with interferometric autocorrelation measurements

M. Fuss, University of Kaiserslautern 13

Optical amplifier characterization (invited)

F.W. Willems, Lucent Technologies, The Netherlands 17

Fiber amplifiers for the second telecommunication window: development and characterisation

Marcello Potenza, Tiziana Tambosso, CSELT 23

All-fiber superfluorescent light source at 1 .06 |im

V. Reichel, A. Ertmer, M. Rothhardt, H.-R. Muller, H. Bartelt, Institut fiir Physikalische

Hochtechnologie e.V. Jena 27

Results of COST 241 EDFA round-robin on gain and noise figure

F.W. Willems, Lucent Technologies, The Netherlands; M. Artiglia, CSELT; C. Larsen,

Lycom; C. Soccolich, Lucent Technologies, USA; B. Heens, FPM 31

A general method for measuring the electric-field-dependent absorption coefficient in quantum

confined structures

J.F. Siliquini, M.G. Xu, J.M. Dell, The University of Western Australia 35

Measuring the reliability of fiber optic components (invited)

L.A. Reith, Bellcore 39

Opto-mechanical method for measurement of ferrule concentricity and roundness error Paul Tovmley-Smith, Costas Saravanos, Siecor Corp 45

Errors due to cormectors in optical fiber power meters

Igor Vayshenker, Xiaoyu Li, Darryl A. Keenan, Thomas R. Scott, NIST 49

MU-type PANDA fiber connector

R. Nagase, S. Mitachi, NTT Opto-electronics Laboratories 53

V
High performance single-mode wavelength independent all-fiber build-out attenuator

Shin-Lo Chia, AMP 57

Nonlinear coefficient of optical fibers at 1550 nm (invited)

Valeria L. da Silva, Y. Liu, A.J. Antos, G.E. Berkey, M.A. Newhouse, Corning 61

Polarization independent nonlinear refractive index measurement in optical fiber

Andrea Melloni, Mario Martinelli, Politecnico di Milano; Andrea Del Core,

Andrea Fellegara, ITALTEL, BUTR 67

Interferometric determination of the nonlinear refractive index n2 of optical fibers

F. Wittl, Techn. Univ. Darmstadt; J. Vobian, G. Herchenroder, W. Dultz,

Deutsche Telekom 71

Inverse method for systematic investigations of linear and nonlinear properties of dispersionshifted and dispersion-compensating fibers

R. Boness, TU Dresden; S. Unger, J. Kirchhof, IPHT Jena; F. Wittl, J. Vobian, FTZ

Darmstadt, K.-F. Klein, FH Giessen-Friedberg 75

Effective area measurement comparison between direct far field scan and variable aperture

method in the far field

Michael J. Hackert, M. Ajaaz Zainul, Coming 79

Characterizing components for high speed data interconnects (invited)

David Smith, Honeywell 83

Modal optical coupling efficiency measurements and modeling for fiber interconnections

J. Sutherland, G. George, J.P. Krusius, Cornell University 89

Dispersion-compensating fibers with high birefringence

J. Vobian, G. Herchenroder, FTZ Darmstadt; K. Mori, IPHT Jena 93

Accurate wavelength calibration for optical spectrum analysers

D. A. Humphreys, National Physical Laboratory 97

Nonlinearity of optical fiber power meters

Igor Vayshenker, NIST; Shao Yang, Ohmeda Medical Systems; Xiaoyu Li,

Thomas R. Scott, NIST 101

Methods of suppressing stimulated Brillouin scattering in optical fibers by manipulation of the

fiber properties (invited)

Clifford Headley, J.B. Clayton, W.A. Reed, Lucent Technologies 105

vi
Local analysis of stimulated Brillouin interaction in installed fiber optics cables

Marc Nikles, Luc Thevenaz, Swiss Federal Institute of Technology; Pascal Salina, Swiss

Telecom PTT; Philippe A. Robert, Swiss Federal Institute of Technology Ill

Non-invasive mapping of dispersion in optical fibers (invited)

R.M. Jopson, Lucent Technologies; M. Eiselt, Heinrich Hertz Institut 115

Determination of zero-dispersion wavelength in optical fiber using four-wave mixing

J.B. Schlager, S.E. Mechels, D.L. Franzen, NIST 121

Accuracy issues in comparisons of time- and frequency-domain polarization mode dispersion

measurements (invited)

P.A. Williams, NIST 125

How accurately can one measure a statistical quantity like Polarization Mode Dispersion?

N. Gisin, B. Gisin, University of Geneva; J.P. Von der Weid, R. Passy, Pontificia

Universidade Catolica do Rio de Janeiro 131

Compensation formula for noise threshold bias of interferometric PMD measurement

B.L. Heffner, Hewlett-Packard Laboratories 135

A practical comparison between two different PMD measurement methods

Bertil Arvidsson, Hans Mickelsson, Ericsson Cables AB; Kim Brising, TELIA 139

Interferometric polarization mode dispersion measurements with femtoseconds sensitivity

Ph. Oberson, K. Julliard, N. Gisin, University of Geneva; R. Passy, J.P. Von der Weid,

Pontificia Universidade Catolica do Rio de Janeiro 143

A novel PMD estimation technique with wavelength stepsize optimization

Osamu Aso, Haruki Ogoshi, Furukawa Electric 147

First report of ITU-T PMD round robin measurements for optical fibers and component

Y. Namihira, KDD R&D Laboratories; ITU-T PMD Round Robin Group 151

TIA round robin for the measurement of PMD

P.A. Williams, NIST 155

Measurement of SOP evolution along a linear birefringent fibre with twist using polarisation

OTDR

R.E. Schuh, A.S. Siddiqui, University of Essex 159

Polarisation dependent loss of concatenated passive optical components

A. Elamari, N. Gisin, University of Geneva; B. Pemy, Swiss Telecom PTT; H. Zbinden,

University of Geneva; Ch. Zimmer, Swiss Telecom PTT 163

vii
Novel method for polarization-maintaining fiber launch

Bernhard Scholl, Jens Rasmussen, Aachen University of Technology 167

Characterization of high-speed and low-loss polymer optical fibers (invited)

Yasuhiro Koike, Eisuke Nihei, Takaaki Ishigure, Keio University 171

A single device for both accurate measurements of and Wq in single-mode fibers

D. Pagnoux, J.-M. Blondy, P. Roy, P. Facq, Institut de Recherche en Communications

Optiques et Microondes 175

Wavelength dependence of correction factor of effective area (Aeff) and mode field diameter

(MFD) in various optical fibers

Yoshinori Namihira, KDD R&D Laboratories 1 79

Fiber cutoff wavelength measurement by modal spectrum modulation

P. Di Bin, C. Simos, D. Pagnoux, P. Faugeras, P. Facq, Institut de Recherche en Communications Optiques et Microondes 183

A simple, fast and accurate measurement of the mode field-radius: The axial gap method

J.L. Auguste, J.-M. Blondy, D. Pagnoux, M. Clapeau, P. Facq, Institut de Recherche en Communications Optiques et Microondes 187

Calibrated test fiber for optical time-domain reflectometers

Duwayne Anderson, Tektronix 191

A method for the localisation and quantization of faults in passive tree-structured optical

networks using the OTDR technique

L. Wuilmart, V. Moeyaert, D. Daniaux, P. Megret, M. Blondel, Faculte Polytechnique de

Mons 195

A coarse ranging measurement method for access systems using passive optical networks

A. Fellegara, ITALTEL 199

Spatial-resolution improvement in coherent optical frequency domain reflectometry for long

optical fibers by reducing frequency-sweep nonlinearity

K. Tsuji, K. Shimizu, T. Horiguchi, Y. Koyamada, NTT Access Network Systems

Laboratories 203

A new wide scan range high resolution optical low coherence reflectometer

Mauro Bottanelli, SIRTI S.p.A 207

Interferometric noise in fiber transmission systems incorporating birefringent substances

Michael A. Bukhshtab, Coming 211

Author Index 215

viii
Characterization of optical components for

D-WDM applications

J. Dupre

Hewlett-Packard Co.

Santa Rosa, CA

1 Introduction

Dense wavelength-division-multiplexed (D-WDM) transmission systems at 1550 nm

have become commonplace for increasing the capacity of point-to-point terrestrial links. In the near future, this technology will be employed as well for undersea systems [1]. In addition, the

all-optical network, providing flexible high capacity through wavelength routing, will soon be a

reality [2]. In both cases, the optical components utilized in these systems require more

sophisticated characterization than those intended for single channel applications. Simply stated,

the wavelength dependencies of the optical parameters of the transmitter, amplifier, passive, and

receiver components, while only of secondary importance in single-channel applications, become

critical to the performance of D-WDM systems. The laser sources comprising the WDM

transmitter must be selected and maintained precisely to the wavelengths required by the

particular channel plan. The passive components such as multiplexers, demultiplexers, isolators,

and couplers must have their parameters, including polarization dependence, measured versus wavelength. The erbium-doped fiber amplifiers (EDFAs) must have their gain and noise figure

characterized for the multi-signal environment in which they are used because the spectral gain

dynamics can have a particularly adverse effect on system performance [3].

This paper addresses measurement issues on two important components of D-WDM

systems: the demultiplexer and the EDFA.

2 Characterizing demultiplexers

Fig. 1 shows a simplified diagram of a D-WDM system for point-to-point transport. The

outputs of N laser sources at the prescribed wavelengths are combined and applied to the

WDM Erbium doped fiber Transmitter amplifiers Demultiplexer

QD-i

Fig. 1 . A D-WDM system for point to point transmission.

Channel N hannel H-1

Fig. 2. Some typical demultiplexer charateristics. In general,

all parameters are a function of input state-of-polarization.

1
transmission fiber. EDFAs are typically used

in booster, line, and preamplifier

configurations. At the far end of the system,

the demultiplexer separates the N channels and

routes each one to an optical receiver. The

perfect demultiplexer would have a rectangular

bandpass at each channel just wide enough to accommodate source laser drift, have minimal

insertion loss, high rejection of other channels,

and no polarization dependence. In practice,

these parameters must all be carefully

characterized as indicated in Fig. 2. There are two basic setups that can be utilized to make

these measurements and each has particular

advantages. The first, as shown in Fig. 3(a), is a narrow linewidth tunable laser in

combination with an optical power meter or

optical spectrum analyzer (OSA). The second,

Fig. 3(b), uses a source of broadband noise

and an OSA. In both configurations, a polarization controller is necessary to probe

the device with all states of polarization.

Which setup to use is determined by the

required amplitude and wavelength

measurement ranges, wavelength resolution,

and measurement speed. The tunable

laser/power meter setup has excellent

wavelength resolution, limited by its

(b)

Fig. 3 Two alternative setups for measuring demultiplexers: (a) narrow linewidth source and power meter or optical spectrum

analyzer (OSA) and (b) broadband noise source and OSA.

Fig. 4. The spectral power density for three types of broadband

noise source. The edge-emitting LED and ASE sources have

sufficient power to adequately characterize demultiplexers.

Measurement

Range

80 dB

with OSA

tunable laser

with power meter

10 nm 1 nm 0.1 nm 0.01 nm 0.001 nm

Wavelength Resolution

Fig. 5. Comparison of measurement methods

2
incremental wavelength linearity to about 0.001 nm. The measurement range is limited by the

tunable laser's spontaneous emission level and the

filter characteristics of the DUT. A narrow

(1-nm) bandpass filter, for example, can be

measured with 40-dB range while a band reject

filter would be limited to about 20-dB. Replacing

the power meter with an OSA or adding a tunable

filter on the source will increase the measurement

range significantly.

E 24 c

£2-2 D

TJC

(Q i,a

fflO 16

-20-dB BW

} ~t i

'

]

-10-dtj uw

-3-dB_BW f1> <

1

1

1 1 =l:iBBW__^

0.1 0.2 OSA Resolution BW, nm

0.5

Fig. 6. Filter bandwidth measurement results using an EELED source and three settings of resolution bandwidth.

The key advantage of the second

alternative ~ the broadband noise source and

OSA ~ is measurement speed which is determined by the OSA's sweep speed. There are several options for the broadband sources shown

in Fig. 4 [4]. An incandescent light bulb, while very broadband, does not have sufficient power

spectral density to measure the out-of-band characteristics of demultiplexer filters. An

edge-emitting LED (EELED) or the amplified spontaneous emission (ASE) from an EDFA with no input signal can provide useful measurement range over the bandwidths required for

demultiplexer characterization. Fig. 5 shows the measurement range and resolution possible

from the alternative test setups. The majority of D-WDM devices can take advantage of the

faster measurement time of broadband noise techniques while for very narrow filters (<0.2-nm

3-dB width) , a setup with a tunable laser is required.

The resolution of a filter measurement made with a broadband noise source and OSA is limited by the OSA's resolution bandwidth. The actual measured result is the convolution of the

OSA's filter characteristic with that of the DUT filter. To gauge the impact of this effect on the

measurement of a typical D-WDM filter, a measurement of a filter's -1, -3, -10, and -20 dB

bandwidths were made with three settings of the OSA's resolution bandwidth (Fig. 6). The

maximum bandwidth error was 8% when the OSA resolution bandwidth was set to about

one-half of the filter's 3-dB bandwidth. This error can be significantly reduced by measuring the

OSA's filter spectral shape with a narrow source and deconvolving it from the measured result.

3 Characterizing optical amplifiers

Measurements on EDFAs in compression are generally made with a single saturating

signal. Gain is simply the output signal power divided by the input signal power. Noise figure is calculated from the amplified spontaneous emission (ASE) measured at the amplifier output

either at or near the signal wavelength. The signal wavelength is tuned across the band of

interest to obtain the noise figure and gain as a function of wavelength. This method is quite

adequate for single wavelength systems but, for WDM applications, the resulting gain and noise

figure results are not an accurate predictor of system performance. In multi-channel systems

(Fig. 7), all of the signals contribute to EDFA gain compression. Each signal amplitude affects

3
the gain at all wavelengths. That is, a change in the ?i, signal power causes a gain change not

only at ^i, but also at X2 through X^. Similarly, ASE changes at all wavelengths for a signal level

change at any wavelength. The problem is compounded in the more common situation for

point-to-point transport in which there are multiple EDFAs or in a wavelength routing situation

where channels are added and dropped. Clearly, characterization techniques are required that can

Po,..Po„

EDFA DEMUX

Figure 7. An EDFA in a multi-signal environment.

Multi-;\. Source

predict and assure system performance in spite of the complex gain dynamics in the EDFA.

There are two ways to approach this characterization. The most direct is to utilize a multi-channel source setup with the exact wavelengths in a particular channel plan. Gain is derived by measuring the signal power with the EDFA bypassed, then with the EDFA in place,

and carefully excluding the effect of source spontaneous emission and amplified spontaneous

emission. Multi-channel noise figures can be obtained by estimating the ASE at each channel

wavelength by interpolating between channels and subtracting the effect for source noise. For narrow channel spacings, this is difficult due to the OSA resolution. Alternatively, time-domain

extinction [5] [6] that takes advantage of the slow time dynamics of the EDFA can be used to estimate the ASE. The cumbersome feature of

this multiple source approach is that a large

number of lasers at specific wavelengths must

be assembled. If characterization is required

for muhiple charmel plans (differing

wavelengths, number of channels, and channel

spacings) then the setup needs to be

reconfigured for each characterization.

Another approach that takes advantage

of the largely homogeneous saturation

property of the EDFA utilizes a reduced

number of saturating lasers and broadband

small-signal probe to measure gain

continuously across a wavelength range [6].

Optical Spectrum Analyzer

( ) 1

1

EDFA

Fig. 8. The direct approach to multi-channel gain and

noise figure measurement.

4
The concept of reducing the number of

saturating lasers is explained in Fig. 9. The

full complement of signals for the

multi-wavelength plan is divided into regions

(a). For each region, a single saturating source

replicates the saturating effect of all the

channels(b). The number of saturating sources

required depends upon the significance of

inhomogeneous effects, namely spectral hole

burning. For erbium doping, at room

temperature, one laser can accurately replicate

multiple chaimels over about 15-nm of

spectral width.

wavelength (b)

Fig. 9. A simplified test scheme results from replacing the n D-WDM channels with a reduced number of sources and using

noise gain profile. A setup for making gain measurements with a broadband noise probe, called noise gain

profile is shown in Fig. 10. An edge-emitting LED (EELED) is combined with a tunable laser

and applied to the test device. In order to obtain the value of the wavelength, Xja,, and optical

power, Ps3„ of the single saturating laser that accurately replicates the saturation of a number of

channels, a recursive algorithm is used that sets the population inversion equal for the two cases

[7]. The algorithm makes an initial estimate of and and measures the noise gain at the

channel wavelengths. Then, a new estimate is made using the following equations until

convergence is achieved.

P

sal P^sal = —^— y 1 G P ^ A,n n n Ajal'Jsai

where P^^, _ Gsat , Xsai are the power, gain, and wavelength

of the single input signal.

P^ , G , P are the power, gain, and wavelength

of the n channel signals

Tunable Laser

a

Optical Spectrum Analyzer ni

)

! ° J

Optical Amplifier

Fig. 1 0. The noise gain profile technique is implemented with one

or more tunable lasers and an EELED probe.

5
An experiment was set up to compare gain measurements for a four channel system using

the direct method with four lasers of Fig. 8 and the noise gain profile method with a single

saturating source of Fig. 10 using the convergence algorithm described above. The results, as shown in Fig. 1 1 , agree to within 0.2 dB which is within the measurement uncertainty of the two

methods.

1520 '1530 1540 1550 1560 1570 1580

Wavelength, nm

Fig. 1 1 . Comparison of direct gain measurement

and noise gain profile.

Acknowledgments

The author would like to thank Douglas Baney, Walter Benger, Jim Stimple, and Joachim

Vobis for contributing material for this paper.

References

[1] Neal Bergano, et o/.,"100 Gb/s Transmission of Twenty 5 Gb/s NRZ Data Channels Over

Transoceanic Distances Using a Gain Flattened Amplifier Chain," ECOC'95, post deadline

paper, 1995.

[2] Paul E. Green, Jr., "Optical Networking Update," Selected Areas in Communications, vol.

14, no. 5, June 1996, pp. 764-769.

[3] B. Clesca, D. Bayart, and J.L. Beylet, "1.5-|a.m Fluoride-Based Fiber Amplifiers for

Wideband Multichannel Transport Networks," Optical Fiber Technology 7, 1995, pp.

135-157.

[4] Dennis Derickson, "Measuring ASE in Fiber Optic Systems," Communication System

Design, January 1996, pp. 36-41

[5] D.M. Baney and J. Dupre, "Pulsed-Source technique for optical amplifier noise figure

measurement," European Conference on Communications, ECOC'92, paper WeP2.11,

Berlin, 1992.

[6] H. Chou and J. Stimple, "Inhomogeneous Gain Saturation of Erbium-Doped Fiber

Amplifiers," Optical Amplifiers and Their Applications Technical Digest, vol. 18, paper

ThEl-1, 1995.

[7] D.M. Baney and J. Stimple, "WDM EDFA Gain Characterization with a Reduced Set of

Saturating Channels," To be published.

6
INVITED PAPER

Design and measurement of optical fibre gratings

Raman Kashyap

BT Laboratories, Martlesham Heath, Ipswich IPS 7RE

United Kingdom

Abstract

Optical fibre gratings for application in telecommunication are now routinely produced. Recent developments have

been in the area of specialist gratings for band-pass, ultra-narrow-band filters with either excellent side mode

suppression or particular phase-characteristics, and also chirped gratings for dispersion management in optical fibre transmission systems. This paper introduces some of these recent developments and address the problems associated

with design, fabrication and measurement of the transfer fimctions for specific applications.

Fibre Bragg gratings were discovered by Ken Hill et o/[l] in 1978 during the course of experiments with a line narrowed argon ion laser operating at 514 rmi and a germania doped silica optical fibre. It was noticed that the transmitted power continued to reduce with time and upon investigation it was realised that the power was being

reflected. The proposed reason for this phenomenon was the inscription of a refractive index Bragg grating within

the core of the fibre. The propagating wave interfered with a counter-propagating wave due to a weak reflection from the far end of the fibre to form a spatially periodic intensity pattern which modified the refractive index of the

core. As more light was reflected, so did the refractive index perturbation grow in concert, with nearly 100% of the

light being reflected after some minutes. This was the beginning of the field of photosensitive optical fibres. The

Bragg grating had the correct period for reflection at the wavelength of the inscribing wavelength only and was

restricted to the visible part of the spectrum. Research was slow to pick up in this field because of difficulties in making measurements and the belief that there was something magical about the original fibre. Stone[2], however,

showed that most germania doped silica fibres demonstrated photosensitivity and following the interesting discoveiy

of second harmonic generation from germania doped silica fibres[3] as well, worldwide activity in optical fibre photosensitive behavior began.

The next major step was taken by Meltz et al[4], who demonstrated that by applying standard holographic

techniques with UV radiation at 248iun, the Bragg wavelength could be shifted to more usefiil longer wavelengths in the near IR at 850mn. By a fiuther modification, gratings at telecommunications wavelengths in the 1500nm

window were reported, opening the area of photosensitive optical fibre Bragg grating devices[5].

Optical Fibre Bragg Gratings:

Bragg rejlection in a waveguide:

Figure 1 shows a refractive index perturbation in a waveguide. A reflection occurs if the period of the perturbation is:

where n,ff is the effective index of the mode in the waveguide, X is the Bragg wavelength, and N is the order of the

interaction[6]. N is one for a first order interaction. For synchronous coupling into the backward propagating radiation modes, satisfying the phase-matching condition. Equation (1) alone, is not sufficient. The second important parameter for efficient interaction is the overlap integral

of the forward tiaveling guided-mode with the backward wave and the perturbation profile. This may be computed

from simple integrals assuming a uniform index-perturbation across the core. For a guided wave reflection, the overlap decreases with either inclination of the Bragg planes for shallow angles, or asymmetry of the perturbation

across the core. On the other hand, counter-propagating radiation mode coupling increases with either asymmetry or

inclination of the Bragg planes[7,8], also for shallow angles. Forward propagating radiation mode coupling requires no asymmetry of the grating across the core; the mode only requires additional momentum equivalent to 27iAn/X,,

where An is the core-cladding index difierence[9].

Introduction:

(1)

7
core

Fibre

A cladding 4-

nil IIIIHIIII IIIIHIIII nil' mil lllll'

Refractive index perturbation

Figure 1. Optical fibre with periodic refractive index modulation in core. Methods ofFibre Grating Inscription:

Figure 2 shows one of the methods of inscribing Bragg gratings in an optical fibre. The interferometer is based

around a beam-splitter[4] or a phase-mask[10,ll,12]. The latter is a phase relief grating made in UV-transmitting

silica which acts as a diffraction grating. In Figure 1, the phase-mask is used as a beam-splitter with two mirrors to recombine the beams at the fibre. The advantage of the phase-mask is that it allows the easy replication of the Bragg

grating at the desired wavelength without re-course to the measurement of the mirror angles. The phase-mask may

be used in contact with the fibre [10], and the UV beam scaimed along the phase-mask to write a long grating[13].

Variations have also been reported[5], including a method to project a magnified amplitude mask on the fibre to write a higher order grating[14]. Another method writes the grating point by point[15]. In this scheme, the index modulation is induced by focusing a UV laser spot at a point in the fibre and then moving the fibre by a period before the second point is written. Although this method allows long gratings to be written, the positional control required for the integrity of first order

grating period, limits this technique to higher orders, or long period gratings[9].

' UVBeam

U U U U u u u u

Phase mask

Rotatable Mirrors

Fibre

Figure 2. Interferometer used for inscription of gratings, with the phase-mask as a beam splitter.

Simple Fibre Bragg Gratings:

A spatially uniform index modulation along the propagation direction produces a reflection spectrum shown in Figure 3, with side-lobes which decrease with de-tuning from the Bragg wavelength. Basically, the optical fibre Bragg grating is a reflecting band-stop filter. This fact alone limits the usefulness of the simple fibre grating as a

device. However, the technology of fibres and Bragg gratings are such that the quality of the perturbation may be

maintained over tens of thousands of periods, producing responses close to those theoretically predicted, so that it has been possible to use the amplitude and phase characteristics of a single grating to shape the spectrum of a

pulse[16]. The side-lobes shown in Figure 3 are undesirable and may be modified by a number of apodisation

schemes[17,18,19]and will be discussed later in the article.

Method ofAnalysis:

Coupled mode theory[20] may be used to calculate the reflectivity of a uniform grating of length, L and dielectric constant peak modulation of as:

8
1540.0 1540.6 1541.2 1541.8

1.0

0.8

0.6

0.2

0.0

-0.2 c -0.4

-0.6

-0.8

-1.0

< 1

<4

44

4: 4

4

< 44

1

i4 4

4

4 4

i 4

; 4 : 1—f-i '

4

4

4 4

4

4

4

4

4

4

T

'

4 4 y\ ^

1 T :\ 4

• 4

4

4

4

4

4 > T

|i : \ 4 U ^ \^

' 4—

'

\ 4

44

44

44>

4

1 4

4 ; 4

4

4 —4

4

4 / AX 4 >

4

4

4

4

44

*

*4

4

44

I 4 -ii-U44

•

4

4

4

44

(4 .

•

i4

y -4

4

44

4

4

4

i4 i 4 44

4

44

4

< 4 4 4 4 44

«

4

4

4

4 4

11 44

4 RX 4 4 4

4

*^ PHASE

1 Wavelength, nm

2.0

1.6

1.2

0.8 o

0.4 .2

at 0.0 «

w

:« -0.4 £

-0.8

-1.2

-1.6

-2.0

Figure 3. Reflection and phase spectrum of a 4mm long Bragg grating with An = 4e-4.

R =

cosh' (OL) + (M / 2)' sinh' (OL)

(2)

where.

K - (3) [2 2 1

''2 K - (M / 2) J is the de-tuning parameter, «o is the average fibre

, 4;r 2A^;r

refractive index and bJi = — is the mismatch between the momentum of the mode and the index

A A

perturbation.

For the special case of zero mismatch, the reflectivity, R at the Bragg wavelength is R = tanh\kL) (4)

Bandwidth, Reflectivity andModulation Index Relationships andMeasurement:

The bandwidUi and the reflectivity are the only parameters required to characterise a uniform grating. The

bandwidth of the grating which is defined as the difference in wavelengths between the first two minima (See Figure

3) is given as AA = -^\{KLy +7r'T (5)

In making measurements or in the design of fibre gratings, the first parameter of interest is the required reflectivity

at the Bragg wavelength. Given this value, the argument of Equation (4) may be calculated. From Equation (5), the length of the grating can be calculated for the required bandwidth. The modulation index is then calculated from

9
Equation (3), with An - As/lriQ . Measurement of a fibre Bragg grating requires a tunable laser or a broadband

ASE source, a circulator or coupler and a spectrum analyser. For long uniform gratings, or gratings of high

reflectivity (AX < 0.2mn or R > 30dB), the transmission spectrum can be difficult to resolve, owing to the resolution

of the spectrum analyser, and require refined spectroscopic techniques for measurement[21].

Numerical Method ofAnalysis:

Gratings of a more complicated nature such as linearly, quadratically chirped or multiple section gratings can be

easily analysed by numerical schemes, for example by Rouard's method[22]. This was first applied to optical multilayer waveguide problems by Weller-Borphy and Hall[23]. This matrix approach allows gratings of arbitrary periods

and refractive index modulation to be accurately modeled, and has been used successfiilly to analyse both linearly

and quadratically chirped gratings[19]. The grating is broken up into M sections, each of distinct constant period.

Each period is also sub-divided into sufficient sections to mimic the contour of the sinusoidal modulation. The total response of the grating is then computed by summing the fields over the length of the Mgratings. The flexibility of

this method makes it highly attractive and reasonably fast for purposes of modeling, producing excellent correlation with experimental results.

Band-Pass Filters: The reflection band-stop filter may be transformed interferometrically into a band-pass filter by using two identical Bragg gratings located on output ports of a 50:50 fibre coupler, equidistant from the coupling region[24]. Balancing

of the interferometer can be performed using UV radiation and the photosensitivity of the core dopant in one arm of

the coupler. Figure 4 shows a schematic of this device.

The important parameter for such a device is the de-tuning from zero-path length difference for identical unapodised

gratings. For a 45dB rejection of the peak, the paths need to be equalised to within ~0.002ABragg [25].

Apodisation offibre Gratings:

Apodisation reduces side-lobes of a uniform grating so that cross-talk can be reduced in fine grain WDM

channels[26]. There are several methods of apodising Bragg gratings. The choice of the type of Banning function is dependent on the application. However, a simple form of apodisation is the cosine apodisation, in which the amplitude of the modulation index of the grating varies as a cosine function around the centre of the grating.

Apodisation requires the Bragg-wavelength to be kept a constant by varying the only the visibility of the fringes at constant UV flux[17,18,19]. An easy and universal scheme for cosine envelope apodisation requires that the fibre be

stretched in opposite direction in a way that the edges of the grating experiences a half-period stretch, while the centre see no change[19]. Periodic stretching while writing the grating automatically gives rise to a cosine envelope

of the visibility function, at constant flux.

Chirped Bragg Gratings:

Ouellette[27] proposed the use of chirped gratings for dispersion compensation. The principle of operation is simple: the point of reflection varies linearly with wavelength, causing a propagation delay which is also a function of

wavelength. Since the propagation delay in a fibre is ~10ns/m in reflection, a compact device is possible; a 100mm

long grating with a chirp of O.lnm is capable of compensating for dispersion of ~700km of standard

telecommunications fibre at a bit-rate of 10Gb/s[28]. However, dispersion compensation is in reflection, requiring a

circulator or a band-pass configuration for the grating to be used[29, 30]. The design of chirped gratings for dispersion compensation requires some form of apodisation[17] if the coherencelength of the signal wavelength is a few percent of the length of the grating[19]. There are several methods of

fabricating a chirped grating. The simplest is to introduce a known strain profile along the length of the grating [3 1,32], or to replicate a grating from a chirped phase-mask[33]. This method is by far the most reproducible

-Bragg

Figure 4. The fibre Bragg grating Michelson interferometer band-pass filter.

10
and mechanically robust method. The step-chirped phase-mask is designed to have n sections, each of a different but

constant period. A relatively few steps are needed for a grating of a given chirp bandwidth, AA^^,,^ (in nm) and

length (in mm) to approximate to a continuously chirped grating[34]. The minimum bandwidth of each section may

be calculated as [3 4] SA^Ux A^^^ (4)

Without apodisation, the reflection spectrum has steep sides and the delay characteristics display a periodic ripple

around an average linear slope; the calculated data of a near continuously chirped grating (200 sections) are shown

in Figure 5 [19]. With a cosine apodisation envelope of the refractive index modulation, both the reflection and the delay characteristics are smoothed, shown in Figure 6.

Miction, May cT (JSnn cMiped ntn i^«lre lOOnm kng: Dn • 6»4

159011 lUOJ M

Wavelength, ran

Figure 6

Figure 5

Figures 5 & 6. Theoretical reflection and delay spectrum of 0.75nm bandwidth, 100mm long step-chirped

Bragg-grating without (Fig. 5) and with apodisation (Fig. 6). The delay and reflection characteristics may be measured by using the scheme shown in Figure 7. A tunable

external-cavity laser is modulated by a external modulator and coupled to a circulator. The grating is spliced to the second port of the circulator, while the reflected signal is detected by a photo-diode and compared with the modulation signal using a vector-voltmeter. As the laser is tuned the reflected signal and delay can be measured

simultaneously. Low coherence reflectometry[35] may also be used to characterise the grating.

Timable

Laser

Modulator

Circulator

Chirped Bragg Grating

^1 I "I

i Photodiode

1 -^—

.

Signal Generator

^ Vector-Voltmeter

Figure 7. Measurement setup for delay and reflectivity measurement

The paper will discuss some of the issues relating to the design, measurement and application of fibre Bragg

gratings.

References:

1 Hill K O, Fujii F, Johnson D C, Kawasaki B S, 'Photosensitivity in optical waveguides: Application to reflection filter

fabrication', Phys. Lett. 32(10), 647, 1978. 2 Stone J, 'Photorefractivity in Ge02 doped sihca fibres', J. ofAppl. Phys. 62, 4371, 1987.

3 Osterberg U and Margulis W, Opts. Lett. 11,516, 1987. 4 Meltz G, Morey W W & Glenn WH, 'Formation of Bragg gratings in optical fibre by transverse holographic method', Opts.

Lett. 14(15), 823. 1989.

11
5 Kashyap R, Armitage J R, Wyatt R, Davey S T, Williams D L, 'All fibre narrowband reflection gratings at ISOOnm', Eectron.

Lett. 26{\\\ 730, 1990. 6 Kashyap R, 'Photosensitive optical fibres: devices and applications'. Optical Fiber Tech. 7(1), 17, 1994. 7 Kashyap R, Wyatt R & Campbell R J, 'Wideband gain flattened erbium doped amplifier using a photosensitive fibre-blazed

grating'. Electron. Lett. 29(2), 154, 1993. 8 Erdogan T & Sipe J E, 'Radiation-mode coupling loss in tilted fibre phase gratings'. Opts. Lett. 20(18), 1838-1840, 1995. 9 Bhatia V & Vengsarkar A, 'Optical fiber long-period grating sensor'. Opt. Lett. 27(9), 692-694, 1996. 10 Hill K O, Malo B, Bilodeau F, Johnson D C, & Albert J, 'Bragg grating fabricated in monomode photosensitive optical fibres by UV exposure through a phase xaas}(.\ Appl. Phys. Lett. 52(10), 1035, 1993.

1 1 Anderson D Z, Mizrahi V, Erdogan T & White E A,, 'Phase mask method for volume manufacturing of fiber phase gratings', Proc. Conf. On Opt. Fiber Comm., OFC '93, Tech. Digest., p68., Postdeadline Paper P16, 1993. 12 Kashyap R Armitage J R, Campbell R J, Williams D L, Maxwell G D, Ainslie B J & Millar C A, 'Light-sensitive optical

fibre and planar waveguides', BT Technol. J. 77(2), 1993. 13 Martin J & Ouellette F, 'Novel writing technique for long highly reflective in fibre', Electron. Lett. 30, 811, 1994

14 Rizvi N H, Gower M C, Goodall F C, Arthur G & Herman P, 'Excimer laser writing of submicrometre period fibre Bragg

gratings using phase-shifting mask projection'. Electron. Lett. 57(1 1), 901-902, 1995. 15 Hill K O, Malo B, Vineberg K A, Bilodeau F & Johnson D C, 'point by point'. Electron. Lett. 26, 1270-72, 1990. 16 Kashyap R, de Lathouwer M, Emplit Ph., Haelterman M, Campbell R J & Armes D J, 'Optical Dark soliton generation using

a fibre Bragg grating'. Topical meeting on Photosensitivity and Quadratic Nonlinearity in Glass Waveguides: Fundamentals and

Applications, OSA Technical Digest No. 301, 1995. 17 Malo B, Theriault S, Johnson D C, Bilodeau F & Hill K O, 'Apodised in-fibre Bragg grating reflectors photoimprinted using a phase mask'. Electron. Lett.31i3), 223, 1995. 18 Cole M J, et ah, 'Moving fibre/phase mask-scanning beam technique for enhanced flexibility in producing fibre gratings with

a uniform phase mask'. Electron. Lett.31{\l), 92, 1995. 19 Kashyap R, Swanton A, & Armes D J, 'A novel technique for apodisation of chirped and unchirped Bragg gratings'. Electron.

Lett. 52(13), June 1996. 20 Yariv A, & Yeh P, In 'Optical waves in crystals ', Ch. 11, Publ. J. Wiley & Sons, 1984. 2 1 Kashyap R McKee P F, Armes D J, Shabeer M & Cotter D, 'Measurement of ultra-steep edge, high rejection fibre Bragg

grating filters', ^/ecfroM.Leff. 57(15), 1282-1283, 1995. 22 Rouard MP, 'Etudes des proprietes optiques des lammes metaliques tres mince', Annal. Phys. II, 7(20), 1937. 23 Weller-Borphy L A & Hall D G, 'Analysis of waveguide gratings: application of Rouard'd method', J. Opt. Soc. Am. B. 2(6), 863-871, 1985. 24 Bilodeau F, Hill K O, Johnson D C & Albert J, 'High-retum-loss narrowband all-fiber bandpass Bragg transmission filter', IEEE. Photon. Technol. Lett. (J(l), 80, 1994. 25 Kashyap R, 'Design of fibre grating band-pass filters'. In preparation.

26 Malo B, Theriault S, Johnson D C, Bilodeau F, Albert J & Hill K O, 'Apodised in-fibre Bragg grating reflectors photoimprinted using a phase mask'. Electron. Lett. 31(3), 223-225, 1995. 27 Ouellette F, 'Dispersion cancellation using linearly chirped Bragg grating filters in optical waveguides'. Opts. Lett. 72(10),

847, 1987. 28 Loh W H, Laming R, Ellis A D & Atkinson D, ' lOGb/s transmission over 700km of standard single mode fibre with 10cm

chirped fibre grating compensator and duobinary transmitter'. Submitted to : Photon. Tech. Lett. 29 Hill K O, Bilodeau F, Malo B, Kitagawa T, Theriault S, Johnson D C & Albert J, 'Aperiodic in-fiber Bragg gratings for

optical fiber dispersion compensation, OFC '94, Postdeadline paper, PD2, pp. 17-20, 1994. 30 Chemikov S V, Taylor J R & Kashyap R, 'All-fiber dispersive transmission filters based on fiber grating reflectors'. Opts.

Lett. 20(14), 1586-1588, 1995.

3 1 Laming R I, Robinson N, Scrivener P L, Zervas MN, Barcelos S, Reekie L & Tucknott J A, 'A dispersion tunable grating in a 10-Gb/s 100-200km step index fibre link', lEE Photon. Technol. Lett. 8(3), 428^30, 1996. 32 Krug P A, Stephens T, Yoffe G, Ouellette F, Hill P & Doshi G, 'Dispersion compensation over 270 km at lOGb/s using an

offset-core chirped Bragg gcatmg". Electron. Lett. 57(13), 1091-1092, 1995. 33 Kashyap R, McKee P F, Campbell R J & Williams D L, 'A novel method of writing photo-induced chirped Bragg gratings in

optical rihies". Electron. Lett. 50(12), 996, 1994. 34 Kashyap R, 'Design of step-chirped gratings'. Submitted to Opts. Comm.

35 Lambert P, Fonjallaz P Y, Limberger H G, Salathe R P, Zimmer Ch & Gilgen H H, 'Bragg grating characterisation by optical

low coherence reflectometry', IEEE Photon. Technol. Lett. 5, 565-567, 1993.

12
Chirp Determination of Fiber Bragg Gratings

with Interferometric Autocorrelation Measurements

M. Fuss, University of Kaiserslautem, Germany

Abstract: The possibility of in-fiber Bragg gratings characterization by analyzing of

the first and second order interferometric autocorrelation functions of ps-pulses fi-om a fiber laser and that ones reflected fi-om the fiber grating has been investigated. The

initial pulses were almost unchirped but the reflected ones obtained a significantly

nonlinear phase shift induced by the fiber grating. Pulses nonlinear phase is dependent on gratings chirp and can be evaluated by means of interferometric autocorrelation measurements.

I. Introduction

In optical fiber transmission systems with high bit rates up to 10 Gbit/s and more dispersion has

to be compensated to minimize pulse broadening for error-free detection. Especially at 1.55^m used

in already-installed fiber links the compensation of fibers chromatic dispersion is of great importance.

Several methods for achieving this are already known but linearly chirped fiber Bragg gratings are a promising method due to their constant dispersion over a bandwidth, sufficient to support

ultrashort pulses [2] [3] [6] [9]. Otherwise chirped fiber gratings specified by customers can find

application as frequency selective components. Chirp determination with the heat-scan technique [7] or by external Bragg reflection [4] have been investigated, however, they require special equipment.

If an interferometric autocorrelation measurement device for ultrashort optical pulses is available it can be used for fiber grating characterization as well.

n. Experimental setup

A measurement configuration consisting of an opto-mechanical autocorrelation apparatus with an electronic control unit and fiinge-counting determination of displacement with digital data

acquisition was constructed [1]. The experimental setup of the autocorrelator is shown in Fig.l.

The light to be measured entered the autocorrelator in a fiber collimator and passed a Michelsoninterferometer with the moving retroreflector mounted on a translatory electrodynamic unit. Two kinds of interferometric signals were measured simultaneously, with a photodetector (PD 1

)

the fijndamental harmonic light interferogram 81(1) as field autocorrelation function Gi(t) and with

a photomultiplier tube (PMT) the Frequency Resolved Second Harmonic Generation (FRSHG)

signal SjCt) containing the intensity autocorrelation GjCx), a portion F,(t) at light frequency and the

second harmonic light interferogram F2('c) at doubled frequency. As nonlinear medium a 3 mm thick

LilOj-crystal was used. It was cut for the center wavelength of 1475nm in order to enable phasematching for 1550nm and 1300nm only by rotating a few degrees left or right on its axis.

To scan the signals with sufficient accuracy the linearly polarized light of a Helium-Neon-laser was directed into the Michelson-interferometer likewise with changing the polarization circularly by

using a quarter wave plate. The two resulting orthogonally polarized interference signals dependent on the movement of one retroreflector were separated with a polarization beamsplitter and detected

with two separated photodetectors. An electronic processing unit generated clock pulses for the

analogous/digital-converter each '/2-AHeNe^316,4nm displacement of the retroreflector.

13
MOVING

RETRO- REFLKCTOR

"7

FIBER CXJLLIMATOR

BEAM- SPUTTKR UlOj FILTER PMT

FIXED RETRO- REFLECTOR

ELECTR. CONTROL

UNIT PC OSCl

Fig. 1: Experimental setup of the coUinear optical interferometric autocorrelator with signal processing electronics

Designating pulse electric field as E(t)=Eo(t)e'"' the measured autocorrelation signals Si(t) and

82(1) can be written normalized as functions of the time delay x between the two beams in the

Michelson-interferometer:

(1) ^j(t) « I

\Eit) +£(/+t)|2 dt oc 1 +Re[G^(x) e^""]

—00

+00

(2) ^2 (t) « 1 1 [E{t) + fl^dto^l + 2 G^ix) + ARe [F^{x) e^"' ] + Re [F^{x) e

^-/"^

Pulse intensity and nonlinear phase particularly can be evaluated without equidistant scanning by

comparing autocorrelation traces with simulated ones assuming physically well founded intensity

profile and nonlinear phase behavior. A simulation program was developed executing this task.

The Fourier transforms of the two with interferometric accuracy digitized autocorrelation signals can be calculated by using Fast Fourier Transformation (FFT). From Si(i) the spectrum |E(a))|^ at the fundamental frequency and from 82(1) the Fourier transforms |I(Cii))|^ at zero frequency and

|U(a))| at the doubled frequency are deduced. Following relations are valid:

(3) |£(a))|2 = ^{G,(t)} = ^{G^{x)] |f/(a))|2 = ^\F^{x)]

Another way of obtaining information about pulses intensity and phase is to reconstruct them

with an iterative loop algorithm using the three Fourier moduli above [5]. The main idea is to approximate the electric field E(t) by replacing the Fourier moduli in frequency domain by calculated ones and applying the relations I(t)=|E(t)|^ and U(t)=E^(t) in time domain.

ni. Method for chirp determination of fiber Bragg gratings

The two autocorrelation signals Si(x) and § (r) of lOps absolutely unchirped pulses were

simulated with 1024 points each. In Fig.2 and Fig. 3 the resulting traces are shown used as references

in comparison with those of the reflected pulses.

The axial distribution of the index modulation affecting an appropriate Bragg reflection condition

regarding wavelength causes different optical pathes for different frequencies if the reflected pulses

14
-50 -<0 -30 -10 -10 0 10 20 30 to 1" Fig. 2: S,(t) of iinchirped pulses

are containing these optical

frequencies. Therefore the

measurement of pulses

nonlinear phase behavior

allows information about fiber

gratings chirp distribution.

Assuming as example a

-50 -to -30 -20 -10 0 10 20 30 to ^ ' Fig. 3: S^Ct) of unchirped pulses linearly chirped fiber grating

extended in z-direction with the

length L and a Bragg wavelength distribution A(z/L) around the center wavelength A^, as follows:

(4) A(—) = A(-) = Ao+AA-=A,+AA — Ar ^ L ^ AT

AT stands for the difference in transit time and is equal to TLn^^^lc. If the optical bandwidth AX of

the pulse signal is matched to the bandwidth of the fiber grating for a fiber length S the dispersion D

can be compensated with AT = DAAS.

Equation (5) for the pulses nonlinear phase <l>(t) is based on the linear chirp of formula (4):

(5)

A/A Cff ^ c AT . AA t . ^ cAAt^

AA Ao Ar 2 a; AT

In accordance to this temporal nonlinear phase shift Si(t) and 82(1) result, shown in Fig.4 and Fig.5.

Comparing the simulated envelopes with the measured signals with attention to the number and

shape of extrem values admits

sufficient conclusion to fiber

gratings chirp.

Parameters for the simulation

of Si(x) and S2(t):

Ao=1.55 \im, AA=0.5 nm,

^ L=10 mm, n,ff=1.5

-50 -to -30 -20 -10 0 10 20 Fig. 4: Simulation of S,(t)

50 -to -30 Fig. 5: Simulation of S^it)

IV. Measurement results

First confirmation of the measurement principle above was carried out with an unchirped fiber

grating. First order autocorrelation function Si(i) of the unchirped fiber laser pulses is shown in Fig.6. The laser signal had to be amplified strongly with an Erbium Doped Fiber Amplifier (EDFA)

before entering the grating, resuhed autocorrelation trace in Fig. 7. Slight Self Phase Modulation

(SPM) causes the trace to appear little compressed. Due to the short pulse duration of

approximately 5ps and gratings narrow bandwidth of about 0.1 nm significant pulse broadening up

to a factor of 10 could be obtained, of both of gratings input and output pulses, shown in Fig.8.

Side maxima illustrate the initial SPM but the basic shape of the envelope has not changed. In this case a further detailed analysis of the measurement results was not necessary.

In a next step a customer specified chirped fiber grating ordered by Deutsche Telekom AG is to be characterized and the results will be presented at the conference.

15
Fig. 6: S,(t) of fiber laser pulses Fig. 7: S,(t) of fiber laser pulses. Fig. 8: S,(t) of reflected pulses by the strongly amplified with EDFA fiber Bragg grating

V. Conclusions

An optical autocorrelator was developed for measuring both intensity and interferometric

autocorrelation flinctions to determine shape and nonlinear phase, especially, of ultrashort pulses. By

comparison or iterative processing of the autocorrelation data aquired with appropriate accuracy,

especially, pulses can be calculated sufficiently. The measurement procedure is suited for

characterization of wavelength distribution in chirped fiber Bragg gratings. In the case of a lineariy

chirped grating the two measurable autocorrelation signals were simulated and significant envelopes

of the traces resulted. An unchirped fiber grating was available up to now in experimental work. The

amplified but nearly unchirped pulses of a fiber laser were fed into the grating and the reflected ones

obtained broadening due to the small bandwidth of the grating but no nonlinear phase change

occured.

VI. Acknoledgements

This work was supported by the Deutsche Telekom AG. The author wish to thank Prof

W.E.Heinlein for constant encouraging and helpfijl discussions.

VII. References

[ 1 ] M. Fuss and W. Heinlein. : Characterization of fibre amplifiers influence on the nonlinear phase of ps-pulses.

Optical Amplifier and Their Applications conference, Monterey, July 10-13, 1 996

[2] K.O. Hill, F. Bilodeau, B. Malo, T. Kitagawa, S. Theriault, D.C. Johnson, J. Albert: Chirped in-fiber Bragg gratings

for compensation of optical-fiber dispersion. Optics Letters, Vol. 19, No. 17, 1994

[3] R. Kashyap, S.V. Chemikov, R.J. Campbell, P.F. McKee, D.L. Williams, J.R. Taylor: Demonstration of dispersion

compensation in all-fibre photoinduced chirped gratings. Pure Applied Optics, No. 4, 1995

[4] P A. Krug, R. Stolte, R. Ulrich: Measurement of index modulation along an optical fiber Bragg grating. Optics Letters, Vol. 20, No. 17, 1995

[5] K. Naganuma, K. Mogi, H. Yamada: General method for ultrashort light pulse chirp measurement. IEEE Journal of Quantum Electronics, vol. 25, no. 6, 1989

[6] F. Oulette, J.-F. Cliche, S. Gagnon: All-Fiber Devices for Chromatic Dispersion Compensation Based on Chirped

Distributed Resonant Coupling. Journal of Lightwave Technology, Vol. 12, No. 10, 1994

[7] S. Sandgren, B. Sahlgren, A. Asseh, W. Margulis, F. Laurell, R. Stubbe, A. Lidgard: Characterisation of Bragg

gratings in fibres with the heat-scan technique. Electronics Letters, Vol. 3 1

, No. 8, 1 995

[8] D. Tavemer, D.J. Richardson, J.-L. Archambault, L. Reekie, P.St.J. Russell, D.N. Payne: Experimental investigation

of picosecond pulse reflection from fiber gratings. Optics Letters, Vol. 20, No. 3, 1 995

[9] J.A.R. Williams, L Bennion, K. Sugden, N.J. Doran: Fibre dispersion compensation using a chirped in-fibre Bragg

grating. Electronics Letters, Vol. 30, No. 1 2, 1 994

16
Optical Amplifier Characterization

F.W. Willems

Lucent Technologies

Network Systems Nederland

P.O. Box 18, 1270 AA Huizen, The Netherlands

Abstract: We present an overview of black-box measurement issues for optical amplifiers for analog and

multi-wavelength characterization. In addition, we

propose figures-of- merit for Erbium doped fibers used

for those applications.

1 Introduction

This paper is divided in two major parts, the first one dealing with definitions and measurement methods for optical amplifiers treated as a black box, the

second one dealing with characterisation of the main

components in this black box: the Erbium doped

fiber.

2 Black-box characterization

2.1 Single wavelength "Digital" Amplifiers

For these "simple" amplifiers, measurement methods

exist for optical output power, gain, noise figure and

reflectance. Basically they can be devided into Optical and Electrical methods. Typical examples for both are the the noise figure measurement by polarization nulling [1] using an Optical Spectrum Analyzer (OSA), and the one by RIN subtraction [2] for the ESA method.

2.2 "Analog" Amplifiers

Non-linear distortion and in particular second order

distortion, is typically associated with amplifiers used

in analog CATV systems. Several sources can be

identified for this second order distortion:

1. Signal induced carrier- density modulation [3,

4]. The gain of an optical amplifier will be modulated by the signal if the product of the carrier lifetime Tc and the signal modulation angular

frequency of subcarrier i, u)i, is below a certain

value, since the second order harmonic distor- tion is given by Eq.(l).

with m the optical modulation index per channel. Pout the optical amplifier's output power

and Psat the optical amplifier's saturation out- put power. For an EDFA we have = 10ms

and Ui = 50 MHz, and therefore u>iTc = 5

10^, which leads to negligible contribution to the second order distortion. However, note

that for a semiconductor amplifier, typically

Tc — 300ps, leading to a value 0.015 for the

lifetime- frequency product, which is low enough

to cause a large amount of distortion. THe carrier lifetime of Praseodymium doped amplifiers

is around 300 /iS which is also long enough to avoid carrier-density modulation. Therefore it

is expected that for fiber amplifier modules, it

is not necessary to measure this effect.

2. Interaction of transmitter wavelength chirp with

the gain-slope at the signal wavelength [5, 6]. A

gain profile around a saturating signal which is not flat, converts the frequency shift by direct modulation of a semiconductor laser to a spuri- ous intensity modulation, causing second order

distortions that can calculated with equation

Eq.(2)

with Suo the frequency chirp amplitude in Hz,

and the gain-slope of the EDFA at the signal wavelength. For directly modulated sys- tems, the chirp amplitude can be calculated

with 6uo - mr]FM{hias - hh), with tjfm the

FM-response of the laser diode, Ibias the bias current and Ith the laser threshold current. This distortion source is always present in principle.

3. Interaction of transmitter wavelength chirp with

dispersion of the Dispersion Compensating Fibre (DCF, if present) [7]. Optical frequency

chirp by direct modulation of the laser current, causes different propagation delays in a dispersive medium for different parts of the information signal, resulting in a distorted receiver in- put signal. If an optical amplifier is equipped

with Dispersion Compensating Fibre (DCF), to

17
compensate for the dispersion in the feeder sec- tion in the network, the second order distortion

will be very huge, if no precautions are taken

to "compensate for the compensation" . The re- sulting second order distortion would be: HDf = [-^hSuoDL^ (3)

with fd the subcarrier frequency for which the

distortion is calculated, D the fibre dispersion,

and L the fiber length. If the dipsersion compensator (DCF, chirped grating) is integrated

in the optical amplifier, large values for CSO

will result if no proper precautions are taken to compensate the compensator.

Interaction of Polarization-Mode-Dispersion

(PMD) and Polarization Dependent Loss (PDL) [8]. Polarization modulation caused by

interaction of laser frequency chirp and PMD of a section just in front of the optical amplifier, is converted into intensity modulation if the amplifier exhibits PDL. Since polarization mode

coupling is a time dependent phenomenon, the

distortion fluctuated in time an only the time

averaged second order distortion can be calculated:

HDpdl Si^oAT < At > (4)

with AT the PDL, and < At > the PMD.

This effect can occur if an element that exhibits PMD (e.g. the Erbium Doped Fiber or the DCF) preceeds an element that has PDL

(e.g. a WDM or an isolator). A separate PMD

and PDL measurement might be needed to re- solve this eff'ect.

Interaction of Self- Phase-Modulation induced

frequency chirp with dispersion of DCF (if present) [7, 9]. SPM results in a frequency chirp

that can interact with the fibre dispersion, re- sulting in second order distortion. The distor- tion is by

2m.^^J^PjjD^ C Aeff

Ct L

with ^2 the fiber's nonlinear index of refrac- tion, Aeff the effective core area, Pp the av- erage optical power entering the fiber, and Of the loss of the fiber. If the optical ampUfier

contains dispersion compensating fibre at a lo- cation such that high optical power is injected

in the DCF, SPM can cause frequency chirp

and the same situation as with chirp-dispersion

interaction exists.

6. Coherent interference of multiple discrete re- flections [10, 11]. Mixing of the signal with

a delayed fraction of itself, distorts the signal

and generates second and third order distortions. The delayed fraction results from multiple reflections (MPI). The second order distor- tion magnitude is given by

'^^J2{z,)l[Mz,) 1 (6) m

with Reff the eff"ective reflection coefficient, Jn the n-th order Bessel function, and Zi —

sin{u)iT/2), and r the cavity roundtrip

time. Occurence if more than one discrete optical reflection exists, that have a distance less than the coherence length of the transmitter

laser.

7. Stimulated Brillouin scattering induced intensity modulation [12] No simple equation to calculate this distortion exists. Occurence:The

linewidth of the transmitter is not well above

the SBS linewidth; DCF is used inside the optical amplifier.

Most of the distortion mechanisms described above

can not be solely attributed to the optical amplifier,

but arise in conjunction with non-idealities from else- where in the system. Ideally an OFA parameter can be measured without having to do a system measurement. Currently, it seems that the only candidate for system independent measurement is the gain slope:

Definition: We define the gain-slope under single

wavelength operation as, "in the presence of a sig- nal at given wavelength and optical input power, the

derivative of the gain of a small probe, versus the

wavelength, at the signal wavelength". The probe

total time averaged power shall be at least 30 dB below the input signal level, to minimize the gain at the

signal wavelength.

"WDM" Amplifiers

The field of application of optical amplifiers in multiwavelength system is becoming larger every day, as is the growth in terminology that is used to address typical WDM parameters. We will therefore first present a set of definitions that follow the, still continuing

discussion, within lEC.

18
2.3.1 Definitions

• Multi-channel gain: The gain for each channel

of a multi-channnel input spectrum for a given

input power configuration, expressed in dB.

• Multi-channel gain flatness: The maximum

channel gain deviation over specified range of channel input power combinations.

• Gain cross saturation: The change in gain of one channel due to a given input signal change on any of the other channels for given input

power combinations, expressed in dB/dB.

• Gain tilt the ration of gain change at any channel to the gcdn change at a selected reference channel due to a change of input condition, ex- pressed in dB/dB.

• Channel addition/removel response: the majcimum change in any remaining channel's output

power following the addition/removel of input

channel(s) after a specified time, expressed in dB.

• Channel noise figure: For each optical channel,

the increase in signal-to-noise ratio at the out- put of an optical detector with unity quantum

efl&ciency that follows an ideal (zero insertion

loss) demultiplexer due to the propagation of a

shot noise limited signal, expressed in dB. It is assumed that all signals are applied at individually specified power levels.

• Channel signal-spontaneous noise : Same as the previous definition, but only includes the

signal-spontaneous beat-noise contribution to channel noise figure, expressed in dB.

2.3.2 Measurement Methods

Within lEC, no measurement methods for multiwavelength application have been defined yet. However the most logical way to analyze WDM behavior

of an optical amplifier is to subject the amplifier to the complete wavelength packet and to measure the

gain and noise figure at each individual wavelength

with one of the methods described above. From a

practiced point of view this has the disadvantage that

multiwavelength transmitter, or at least a number of tunable DFB lasers must be available, which is very

costly.

Considering that the spectral gain behaviour in a homogeneously broadened medium is determined solely

by the average Erbium ion inversion, one can try

to simulate the amplifier's response on a multiwavelength signal, with a single wavelength that generates the same inversion level. The gain spectrum

of a broadband light source consequently gives the

multi wavelength gain spectrum [13]. It has been shown, however, that even at room temperature spectral hole burning can occur [14]. In this case the method above can not be used. However,

currently it seems that errors in a measurement that

uses the noise gain profile method for an amplifier

that exhibits room temperature spectral hole burning, do not exceed 0.2 dB.

3 Erbium Doped Fiber Characterization

3.1 FOIVE for analog applications

The CSO (Composite Second Order) distortion

caused by the interaction of laser chirp and gain-slope

of an EDFA is given by the product of the number of distortion products that generate second order distortions, Ncso (determined by frequency channel allo- cation for a particular CATV system), and harmonic

distortion from one of those products. We found, us- ing Giles modeling [15], that the total second order

harmonic distortion caused by gain-slope is given by

CSOgsl,edfa = Ncso

da{u)

,

—-- _ dg*^ —--

(7)

with /? the laser chirp in Hz/mA, L the fibre length, a the absorption in Np/m, u the optical frequency, g* the stimulated emission in Np/m and n2{z) the

averaged (time and location z) Erbium ion inversion

level. This equation tells us that the distortion can be

minimized by three parameters: the Erbium doped

fiber length, the average inversion, and the emission

and absorption profiles. These parameters also deter- mine the noise figure and output power, and therefore a balance must be found between them to make

the EDFA fulfill the specification. Note that the har- monic distortion vanishes if we zero the term between

square brackets in eq.8, i.e. if

dg*{iy) - 1 da{v)

dv

(8)

Therefore, each wavelength requires its particular in- version to have zero gain-slope. This inversion is given by:

^2(2) = 7:toT (9)

19
For the best LOFA EDF, n{X)D2=o does not depend

on wavelength and is as small as possible in the analog

wavelength range (1545-1560 nm). Therefore, we in- troduce the following EDF figure of merit A to quickly

assess the suitability for LOFA application:

A= (10)

must still be made to base the FOM to be defined

lateron on the relative F figure or on absolute gain

and ripple figures.

r = AG

(16)

[a{X) - ao)]in2{z) - 1) + n2{z)[g*{X} - g*]

{ao +g^)n2{z) - (ao + 6)

where n(A)^2=o the maximum and n(A)g^2=o

the minimum value of n{X)D2^o in the 1545-1560 nm

wavelength range.

3.2 FOM for WDM applications

According to Giles modeling we have,

dP{z,X)

dz = ^{z,X)P{z,X)

where

_ 1

AG

We will use equation 15 to analyze how well different

fibers perform with respect to the two criteria mentioned in the introduction.

A flat noise-gain profile for a certain input

power/gain requirement. The required gain, which we

will denote by Go, is determined by the fiber length

L and the inversion n2 (see Eq.( 13)), the profile AG

can be tuned by n2(see Eq.( 15)). Therefore we must

find for a certain gain, the value for n2 that gives

the best gain-flatness. With this value we can cal- culate the required fiber length with Eq.( 14)), and we "only" have to provide enough pump power to generate the required inversion level. Easy adjustment for new input power arrangement by

changing the pump power Consider next a change in total signal input power. If we want that the gain is not affected by this, we must adjust the pump power

such that the inversion level does not change. Since

the length L is fixed, also the gain profile does not change if we succeed in providing the right amount

of pump power.

Now, the lower the required n2 is, as determined in the previous step, the "easier" a pump power adjustment will lead to a new optimum situation.

Therefore, to summarize, for certain Erbium doped

fiber, we have to look for the inversion level that leads

to the flattest noise-gain profile. The lower this in- version level is, the higher the change is that we can

design a WDM EDFA that can cope with input power

changes.

A good WDM fiber has a flat noise-gain profile over a large range of wavelength for low average inversion. A figure of merit incorporates those parameters. Therefore, we propose the following FOM for WDM applications:

FOMwm = (aA(1 (17)

with AA the wavelength range for which the gain pro-

[a{X) - ao)]{n2{z) - 1) + n2{z)[g*{X) - ^g] file stays within ±0.5dB, and n2{z)^^, the inversion

7(A, z) = [a(A) + g*{X)] n2{z) - [a{X) + 6{X)] (12)

with a the absorption in Np/m, g* the stimulated

emission in Np/m, 77.2 the time averaged inversion

coefficient, and 8 the background loss (Rayleigh

backscatter and impurity absorption in Np/m). If we define the time and location averaged inversion

2; /g n2{z) n2{z)dz, with L the Erbium doped

fiber length, we arrive, expressing all parameters in dB, at the following expression for the gain G

G(A) = L[{a{X) + g*{X))n2{z) - (a(A) + 6)] (13)

The required length for a certain gain and inversion

is given by

L = G

{a + g*)n2{z) - {a + 6) (14)

The difference in required gain at wavelength Aq and

the gain at any wavelength A, Go — G(A), can be cal- culated with equation 13 combined with equation 14:

{ao + 50)^2(2) - (ao + S) (15)

Of course, one can rewrite Eq.( 15) to resemble the

gain-ripple relative to a certain gain-level. The choice

level that leads to a maximum gain flatness. We will show during the presentation data for FOM

analog and WDM and indicate the correlation be- tween the two.

20
4 Conclusions

Measurement methods for single channel "digital"

optical amplifiers are close to standardization. Both

the ESA and OSA methods can be used to get comparable data in principle. Currently under development

axe measurement methods that address the specific

parameters for analog and multi-wavelength digital

OFAs. Figures of merit for "analog" and "WDM"

Erbium doped fibre, have been presented, and a cor- relation between the two has been identified.

References

[1] J. Aspell, J. F. Federici, B. M. Nyman, D. L. Wilson,

and D. S. Shenk, "Accurate noise figure measure- ments of erbium doped fiber amplifiers in saturation

conditions," in Digest of Conference on Optical Fiber

Communications '92, San Jose, pp. 189-190, Feb.

1992. paper ThA4.

[2] F. W. Willems, J. C. van der Plaats, C. Hentschel,

and E. Leckel, "Optical amplifier noise figure de- termination by signal RIN subtraction," in Symposium on optical fiber measurements, Boulder, pp. 7—

9, Sept. 1994.

[3] A. A. M. Saleh, T. D. Darcie, and R. M. Jopson, "Nonlinear distortion due to optical amplifiers

in subcarrier-multiplexed lightwave commimications

systems," lEE Electronics Letters, vol. 25, pp. 79-80,

Jan. 1989.

[4] L. K. Chen, K. Y. Lau, and D. R. Huber, "Fundamental distortion charactersitics of the Erbium fiber amplifier," in Digest of Conference on Optical Fiber

Communications '95, San Diego, Feb. 1995. paper WL6.

[5] C. Y. Kuo and E. E. Bergmann, "Erbium-doped fiber ampHfier second order distortion in analog finks and

electronic compensation," IEEE Photonics Technol- ogy Letters, vol. 3, pp. 829-831, Sept. 1991.

[6] S. L. Hansen, P. Thorsen, K. Dybdal, and S. B. Andreasen, "Gain tilt of Erbium-doped fiber amplifiers

due to signal-induced inversion locking," IEEE Photonics Technology Letters, vol. 4, pp. 409-411, Apr.

1993.

[7] M. R. Phillips, T. E. Darcie, D. Marcuse, G. E. Bodeep, and N. J. Frigo, "Nonlinear Distortion Generated by Dispersive Transmission of Chirped Intensity-Modulated Signals," IEEE Photonics Technology Letters, vol. 3, pp. 481-483, May

1991.

[8] C. D. Poole and T. E. Darcie, "Distorsion related to Polarization-Mode Dispersion in Analog Lightwave

Systems," IEEE Journal of Lightwave Technology,

vol. 11, pp. 1749-1759, Nov. 1993.

[9] C. Desem, "Composite second order distortion due

to self-phase modulation in externally modulated

optical AM-SCM systems operating at 1550 nm,"

lEE Electronics Letters, vol. 30, pp. 2055-2056, Nov.

1994.

[10] F. W. Willems and A. M. J. Koonen, "The influence

of multiple path interference in optical fibre amplifiers on RIN, CSO and CTB in AM-SCM lightwave

systems," in Proc. of European Conference on Optical Communications 1992, Berlin, Sept. 1992. paper

TuA4.5.

[11] J. H. Angenent, I. P. D. Ubbens, and P. J. de Waard,

"Distortion of a multicarrier signal due to optical re- flections," in Proc. of European Conference on Optical Communications 1991, Pans, pp. 569—572, Sept.

1991. paper WeC8-4.

[12] H. Yoshinaga, "Influence of stimulated Brillouin scattering on nonlinear distortion in SCM video

transmission," lEE Electronics Letters, vol. 29, pp. 1707-1708, Sept. 1993.

[13] H. Chou and J. Stimple, "Inhomogeneous gain satu- ration of Erbium doped feber amphfiers," in Optical

Amplifiers and Their Applications, OAA '95, Davos,

pp. 92-95, June 1995. Paper ThEl.

[14] A. K. Srivastava, J. L. Zyskind, J. W. Sulhoff, and

J. D. Evankow, "Room temperature spectral hole burning in Erbium-doped fiber amphfiers," in Digest

of Conference on Optical Fiber Communications '96, San Jose, pp. 33—34, Feb. 1996. paper TuG7.

[15] C. R. Giles and E. Desurvire, "Modeling ErbiumDoped Fibre Amplfiers," IEEE Journal of Lightwave

Technology, vol. 9, pp. 271-283, Feb. 1991.

21

FIBER AMPLIFIERS FOR THE SECOND TELECOMMUNICATION WINDOW:

DEVELOPMENT AND CHARACTERISATION

Marcello POTENZA, Tiziana TAMBOSSO

CSELT - Via G. Reiss Romoli, 278, 1-10148 Torino (ITALY)

Abstract Optical amplification in the 1300 nm range is attracting much interest in the field of Telecommunications today. The present work discusses the realization and the characterization of a 1300 nm amplifier based on a praseodymium doped fluoride active fibre.

1. Introduction The evolution of transport and access networks would be positively

influenced by the development of Optical Amplifiers (OAs) for the second telecom window.

Even though attenuation is higher at 1300 nm, OAs for the second window would allow the upgrading of presently installed systems on G.652 optical fibers as an alternative to the use of third window systems with Erbium-Doped Fiber Amplifiers (EDFAs), which require narrow

linewidth optical sources and dispersion compensation techniques. The avai'ability of

efficient 1300 nm OAs would also be interesting for subcarrier analogue and d gital video

signal systems and other broadband services in the future assessment of the access network.

Semiconductor optical amplifiers (SOAs) are available since several years for application

around 1300 nm. Both bulk and Multi-Quantum Well active layer devices with very good

characteristics have been developed [1, 2]. However, the high coupling losses to the transmission line affecting SOAs and their marked nonlinearities have prevented a widespread use of these devices to compensate for propagation or branching losses in optical

networks. On the other hand, neodymium-doped fiber OAs have been investigated since ten years, but up to now they have shown rather poor gain characteristics. Since 1991, a new type

of OA for the 1300 nm spectral region, i.e., the Praseodymium-Doped Fluoride Fiber

Amplifier (PDFFA) [3, 4, 5] has been studied and developed. It has been demonstrated to reach

small-signal gains of 30 dB with noise figures around 6-7 dB: a first generation of these

devices is already on the market.

r Pr doped ZBLANBbre

INPUT OUTPUT

IwdmT!-^<-

ISOI BF BF B02

2x2

X = fusion splice t= special mechanical splice ISF = lourmediate Saica Fibre X _l

A.

SOURCE

available for a second pump laser

Fig. 1 - Set-up of the Praseodymium-Doped Fluoride Fiber Amplifier.

This paper briefly discusses the development of a PDFFA and the system

characterization of the amplifier in terms of a transmission experiment and of an evaluation

of possible impairments due to non linear effects. Results obtained so far confirm that such

OAs can represent a viable solution to upgrade presently installed G.652 optical links.

2. Realization of a second window optical fiber amplifier Our laboratories

have set a research project for the investigation and development of optical amplifiers in the second telecommunication window. In the framework of this project, called SWORD (Second

23
Window Optical Repeater Devices), a first result has been the realization of a portable, racksized PDFFA, the set-up of which is shown in Fig. 1. As described elsewhere [5], the PDFFA

29

27

m 25

23

c 2 1

"re O 1 9

1 7

1 5

50 -40 -30 -20 -10

Input Power (dBm)

Fig. 2 - PDFFA gain as a function of the input signal power. From top to bottom, the curves refer to the following signal wavelengths: 1300, 1310, 1290, 1320 and 1280 nm.

uses a high-NA (0.4), small core (<2 |j.m diameter) fluoride active fiber. Silica glasses cannot

be used as Pr^"*" hosts because of their high mean phonon energy which makes non radiative

relaxation dominant over radiative processes. The optimum cut-off wavelength is around

840 nm and the Pr-''*"- ion concentration ranges between 500 and 1000 ppmw in order to minimize the incidence of cooperative relaxation events. The jointing of the fluoride fiber to standard fibers is critical as the difference in the glass phase transition temperature between

9,5

9

S 8.5 ^ 8

7,5

7

6,5

6

V)oz

>

50 -40 -30 -20

Input Power (dBm)

1 0

Fig. 3- PDFFA noise as a function of the input signal power. From top to bottom, the traces refer to the following signal wavelengths: 1320, 1310, 1300, 1280, and 1290 nm.

the two matrices makes fusion splicing not viable. The problem can be approached using an

Intermediate Silica Fibre (ISF) having geometrical and optical parameters as close as possible

to those of the fluoride active fibre [5]. The mechanical coupling ISF/active fibre is made by

gluing the two fibers with UV-curing resin into a quartz capillary tube. Angle-cleaved cuts (around 25°) are needed in order to prevent the onset of laser oscillations, since internal

24
gains higher than 36 dB are easily obtainable: robustness and reliability of such a joint are

satisfactory. Finally the joint ISF/standard fibre is optimized by means of the multiple arc technique. Unfortunately, the efficiency of PDFFAs is much lower than that of EDFAs: a 30 dB gain amplifier requires 300 mW pump power around 1017 nm, which is the Pr^^-ion

peak absorption wavelength. However, such a power cannot be easily reached with presently

available semiconductor lasers. For this reason a detuned pumping with rack-sized Nd:YLF

lasers at 1047 nm wavelength is usually preferred, but the pump requirements at such a wavelength are further increased at 600 mW levels. A bidirectional pumping scheme with a Nd:YLF laser was adopted, as shown in Fig. 1.

1 ,00E-04

1 ,00E-05

1 ,00E-06

1 ,00E-07

m 1,00E-08

CD

1 ,00E-09

1 ,00E-10

1 ,00E-1 1

1 ,00E-1 2

PD -F/ )0 cm 52

L 1i

ba( ;k-t o-b acl

1 -36 -34 -32

Received Power (dBm)

-30

Fig. 4 - BER performance of the PDFFA booster over 100 km G. 652 fiber transmission.

Best results obtained so far (A=1300 nm) are insertion loss 11.4 dB, net gain 27.6 dB (-30 dBm

signal input) and saturation power 14 dBm (-6 dBm input). The noise figure is around 7 dB,

2 dB of which are due to input coupling losses: these could be greatly reduced using an ISF more similar to the fluoride fiber (the presently used ISF has NA=0.32). Figs. 2 and 3 show

gain and noise versus input power for the SWORD 2.2g PDFFA amplifier, at different signal

wavelengths. Using two pump sources, the device reaches an output power of nearly 20 dBm.

3. Characterisation of the PDFFA The system characterization of the PDFFA has

been made in a digital transmission experiment. A standard DFB laser emitting at 1311 nm

was directly modulated at 2.5 Gbit/s with a 2^'-l PRBS. This source injected a -2.5 dBm

signal into the PDFFA, corresponding to a 13 dBm signal output power. A 100 km span of

G.652 fiber was used for the transmission experiment, with the SWORD 2.2g amplifier

boosting the source signal. Light was detected using a p-i-n FET receiver with a -32.5 dBm

sensitivity at a 10'^ BER in back-to-back configuration, see Fig. 4. When using the booster

followed by a variable optical attenuator simulating the transmission span, no penalty is observed with respect to the direct source-receiver coupling. Fig. 4 shows that in the actual experiment, the BER penalty is limited to 0.5 dB with respect to back-to-back configuration.

This result confirms the viability of using such PDFFAs for the upgrade of already installed G.652 optical links. The capability of reducing coupling losses at the input of the PDFFA

below the present 2 dB value, also shows that this result can still be improved.

25
When transmitting high power signals over long distances, it is possible that nonlinear effects could produce transmission impairments. That is the reason why work is in progress to compare non linearities in the second and third window. Because of its lowest

threshold, stimulated Brillouin scattering was investigated first. A threshold of 8-9 mW was

O

u

oo

ocO

E

(0c

Incident Power (dBm)

Fig. 5 - Brillouin effect threshold.

found for both an external cavity tunable semiconductor laser and the DFB laser previously

used for the transmission experiment. These figures are quite similar to those reported at 1550 nm. Further measurements are in progress concerning the efficiency of four-wavemixing in the second window, the results of which will also be presented at the symposium.

4. Conclusions Until few years ago, only SOAs were available for 1300 nm telecom

applications. Their advantages were known and employed in logical and in signal elaboration

applications, and for optical loss compensation in integrated optic circuits. On the other

hand, their strong nonlinearities affected applications in line amplification and multichannel

systems. Now, second window OFAs seem to have reached a certain maturity and a first generation of such devices is already available on the market. The development of sufficiently

intense semiconductor lasers for peak absorption LD pumping of PDFFAs would be a key topic

for the realization of a second generation of such amplifiers, able to upgrade presently

installed G.652 optical links and to promote a balanced evolution of the optical network.

Acknowledgement We thank Dr. M. Artiglia and Dr. P. Di Vita for useful discussions

during the preparation of this work.

References

[1] S. Tsuji et al., "High-gain polarization-independent semiconductor optical amplifier with a large optical cavity and angled buried facets". Opt. Ampl. & Their Appl., 1990, PDP-5.

[2] L.F. Tiemeijer et al., "Progress in 1.3 (im Polarization Insensitive Multiple Quantum Well

Laser Amplifiers", Opt. Ampl. & Their Appl, Breckenridge, (CO, USA), 1994, WDl.

[3] Y. Miyajima, T. Sugawa, Y. Fukasaku, "38.2 dB amplification at 1.31 ^im and possibility of

0.98 |im pumping in Pr^^-doped fluoride fiber". Electron. Lett. Vol. 27, p. 1706 (1991).

[4] T.J. Whitley et al., "High output power from an efficient Praseodymium-doped fluoride

fiber amplifier", IEEE Photon. Technol. Lett. Vol. 4, p. 401 (1993).

[5] M. Potenza, B. Sordo, T. Tambosso, "Second Window Optical Amplifiers: Comparison of

Different Pumping Wavelengths and fibers", Opt. Fiber Technol. Vol. 1, pagg. 322-326 (1995).

26
All-Fiber Superfluorescent Light Source at 1.06 fim

V. Reichel, A. Ertmer, M. Rothhardt, H. -R. Muller, H. Bartelt

Institutfur Physikalische Hochtechnologie e. V. Jena

Helmholtzweg 4, 07743 Jena, Germany

Introduction

For different applications in the large field of measurement technique it would be very

useful to have powerful and broadband light sources because the available white light sources often don't have enough power to solve the measurement problem. In some

wavelength regions (0.85 ^m, 1.3 fim) such sources already exist. Mostly these are semiconductor laser diodes which work below their laser threshold or superluminescence

diodes. In the 1.55^m region a very simple solution was found some years ago. Using an Er-doped fiber amplifier without any input signal (noise source) it is very easy to create a broadband and powerful light source for this wavelength. The same is possible for the

Nd:YAG-region of 1.06 nm by using Nd-doped fibers as was experimentally shown by Liu

etal. [1]. In context with the development of silica based, Nd-doped double-clad fibers for the

application in high-power fiber lasers [2] we have investigated the spectral behaviour of the

emission in the wavelength region around 1.06 jim in the nonresonant and the partly

resonant case, respectively.

As an result we present a light source based on a specially designed rare-earth doped fiber

with a spectral bandwidth of more than 5 nm and a peak power of more than 1 mW in the

nonresonant single-pass gain regime.

27
Setup

The principal approach for the investigation of the superfluorescence light source can be

seen in figure 1

For pumping the Nd-doped fiber we used a Peltier-cooled (TEC) laser diode with FCconnector at 810 nm emission wavelength and a maximum output power at the end of a multimode fiber pigtail (length: 2.5 m; diameter cladding/core: 140 /125 ^m) of 800 mW

from SIEMENS AG (Typ 487406).

The Nd-doped double-clad silica fiber under test was manufactured in our laboratory. The

first cladding surrounding the Nd-doped monomode core is a pure silica core with a

diameter of 130 /xm and a numerical aperture (NA) of 0.38. The second cladding is a

silicon rubber with a low refractive index. The Nd-concentration amounts to 1300 ppm and

the fluorescence lifetime 430 /zs. The pump light is coupled into the active fiber by a fiber-fiber coupling with the help of a

3-axes fiber launch system.

The spectral behaviour was measured by a PC-controlled optical spectrum analyser (OSA)

in dependence on the fiberlength, the reflectivity at the fiber endfaces and the pump

power.

fiber pigtail Nd-doped fiber

OSA

3D fiber-fiber

coupling

power supply

+TEC PC

Figure 1: Setup scheme ofthe superfluorescence light source

28
Results

First of all we measured the superfluorescence spectra (ASE) in forward and backward

direction in the single-pass gain modus depending on the pump power and compared these

results with our numerical simulations using the relaxation method [3].

Figure 2 shows a typical spectrum of the superfluorescence light source. In this case the

peak power is around 2 mW and the spectral bandwidth (FWHM) is 6 nm. The spectral

resolution of the spectrum analyzer is 0.5 nm.

1,0

0,8

[au] 0,6

0,4

outpi 0,2

0,0

1 I 1020 1040 1060 1080 1100 1120 1140 wavelength [nm]

Figure 2: Spectral behaviour ofa superfluorescence light source at 1.06 (xm; fiber

length : 70 m

In addition, we measured the dependence of the output power and the spectral bandwidth on the fiber length. We investigated different lengths between 10 and 70 m.

In figure 3 we demonstrate the dependence of the spectral bandwidth on the pump power in

the nonresonant case. It can be easily seen that the bandwidth drops down with increasing

pump power. When the laser threshold is reached the source begins to lase and the

broadband continuous spectrum changes to a typical mode spectrum of a laser. Hence, the

spectral bandwidth decreases to the bandwidth of the single modes.

Following these measurements in the single-path gain regime the reflectivity at the

outcoupling end of the fiber was increased step by step. At first we cut the fiber endfaces

without any angles (reflectivity 4%) and then we used dichroic mirrors with a high

reflectivity for the pump light and a partially reflecting (24-76%) for the signal wavelength.

This way we approached a double-pass gain device.

The output power increased step by step to several 10 mW, whereas the spectral bandwidth

decreases slowly. Detailed results will be demonstrated at the conference.

In a first application we used the superfluorescence light source for the online control and

characterisation during the manufacturing of fiber Bragg gratings at 1.06 ^m.

29
E

c

12-

10-

14 -

8-

6

0 100 200 300 400 500 600 700 pump power [mW]

Figure 3: Dependence of the spectral bandwidth on the pump power

Conclusion

In conclusion we developed a powerful and broadband superfluorescence light source in the

very interesting wavelength region of 1.06 /xm using a Nd-doped double-clad silica fiber.

We presented experimental results on the behaviour of the light source, characterizing the

linewidth and output power with variations in pump power, fiber length and reflectivity at one fiber end.

In principle, this technique can be used for the telecommunication wavelengths too by

substituting the rare earth element Neodymium by Praseodymium and Erbium,

respectively.

We have shown that this light source is a useful instrument for the online control during the

manufacturing of fiber Bragg gratings.

Some further applications for this kind of light source exist in the field of measurement

technique and especially in the field of fiber sensors (e.g. gyroscopes).

References

[1] K. Liu et al.: " 10 mW superfluorescent single-mode fibre source at 1060 nm",

Electron. Lett. 23 (1987), pp. 1320-21

[2] H. Zellmer et al.: " High power cw neodymium-doped fiber laser operating at 9.2W with high beam quality". Optics Lett., 20 (1995), pp. 578-80

[3] V. Reichel et al.: " Spectral Behaviour of Nd-doped Double-Clad Fibers", will

be presented at CLEO/EUROPE 1996, Hamburg

30
Results of COST 241

EDFA Round-Robin on Gain and Noise Figure

F.W. Willems

Lucent Technologies

Network Systems Nederland BV

P.O. Box 18, 1270 AA HUIZEN, The Netherlands

M. Aniglia, CSELT, Italy

C. Larsen, Lycom, Denmark

C. Soccolich, Lucent Technologies, USA

B. Heens, FPM, Belgium

Abstract: We present gain and noise figure measurement results from an EDFA round-robin

using the OSA methods. This round-robin was performed in the framework of COST 241.

1. Introduction

COST 241, "Characterization of Advanced Fibers for the new Photonic Networks", working

group 3, study group 2,"Fibre Amplifier Engineering", is currendy performing an EDFA

round-robin among the participants of this group. After some initial evaluation of the

inticacies of such a round-robin [1], we started to use an AMOCO 17 dBm optical amplifier

for the round-robin. This device, besides being extremely sturdy, shows a rather peculiar

wavelength characteristic and is therefore very well suited to evaluate the various

measurement methods for gain and noise figure, as they are currently evaluated. In the present

round-robin, we have restricted ourselves to the "optical" measurement methods.

2. Measurement methods

Along with the Round-Robin module, copies are supplied of the lEC documents that describe

the measurement methods for gain and noise figure. In particular: lEC SC 86CAVG3

documents 1290-1-1 (Gain by OSA) [2] and 1290-3-2, (Noise by OSA) [3] are of relevance

for this round-robin. Next we will shortly describe these measurement methods.

Gain by OSA; the optical source can be either di fixed wavelength optical source (DFB, DBR,

External Cavity Laser or a LED with optical filter) or a wavelength tunable optical source. In

the calibration step, the optical powers entering the Device Under Test (DUT) are calibrated

against an optical power meter. Read-outs from the Optical Spectrum Analyzer (OSA) are aligned with these data. The OSA can now be used for calibrated optical power

measurements. The gain is defined as output power minus input power, with corrections for

the Amplified Spontaneous Emission (ASE).

Noise Figure by OSA: The noise figure, NF, basically, is an "electrical" parameter, since it is by definition the input Signal-to-Noise Ratio (SNR), divided by the output SNR, assuming a

shot noise limited signal source. However, theoretically, it can be shown that, at least in the

31
small signal regime, the NF can also be determined by optical means, by measuring the ASElevel in a certain optical bandwidth B, and by measuring the gain G. The signal-spontaneous

beatnoise contribution to the NF can then be calculated by NF=ASE/(hvGB).

To determine the ASE level, currently three methods exist: • interpolation: the ASE curve around the signal wavelength is measured and a curve fit is applied to the data; the value of the fit at the signal's wavelength is taken as the ASE value

for the NF calculation. The disadvantage of this method is that the measured curve is a

superposition of the ASE and the amplified Source Spontaneous Emission (ASSE),

therefore one has to correct for the ASSE.

• polarization nulling: the polarized signal component is filtered from the total optical signal

from the DUT, leaving the ASE perpendicular to the signal and its ASSE to be measured

by the OSA. The disadvantage of this method is that it measures not the ASE component

giving rise to s-sp beatnoise. Polarization Hole Burning effects are not taken into account

by this method.

• pulse recovery: the input optical signal is temporarily shut down. The total ASE is measured at a certain defined point in time and the ASE in saturation is calculated using

assumed ASE recovery.

Five participants, A-E, were included till now in the round-robin. All used the Polarization

Nulling Technique. One lab, E, also used the interpolation and the pulse recovery technique.

3. Measurement results

The figures below show the gain and noise figure at 1550 and 1535 nm respectively, as measured by the particpant in the round-robin. First let us compare the gain data as measured

with the polarization nulling technique. The spread at 1535 nm is around 3 dB and at 1550

nm around 2 dB. However note the inconsistancy in the data for the two wavelengths from

participant E. Also, a large difference exist between the data from different methods as used

by participant E. The interpolation method gives an overestimate and the pulse method gives an underestimate, however this tendency is again largest at 1535 nm.

The spread in the noise figure data is, as expected, even larger than the spread in the gain

data. However, the spread is now also a function of input power. It is interesting to note that

the spread is nearly equal for the two wavelengths. One would have expected that the noise

figure measurement at 1550 nm would give less room for error, since the ASE spectrum is flatter in this wavelength region. Large differences occur especially at the high signal input

powers, since the results are in this case heavily dependent on the quality of the signal

extinction. The large difference between the 1535 noise figure data below -25 dBm input

power of participant D and the rest of the data can be attributed to the fact that also D's gain

data is high compared to the rest of the field, possibly because of a flaw in the ASE

subtraction procedure. Note the deviation of the interpolation and polarization nulling data

from participant E from the rest of the data, including E's pulse data, which fits quite well

with the rest of the field.

32
Since the noise figure by optical method is determined by the quotient of ASE level and gain, a systematic error in the optical power measurement is hidden in the noise figure data, if this

is the only error source. This is shown nicely in the noise figure data for 1550 nm, where the

noise figures as determined by A-D cluster very well, while the gain data show a large spread.

GAIN @ 1650 nm vs INPUT POWER

32.00

GAIN @ 1535 nm vs INPUT POWER

33
noise figure @ 1 535 nm vs input power

14.00

6.00

-40 -30 -20 -10 0

Input Power (dBm)

4. Discussion and conclusions

Five participant used the polarization nulling method to measure gain and noise figure of an

optical fiber amplifier, in a first international round-robin organised in the COST 241

framework. The spread in the measurement data is more than a few dBs. Possible causes for

this are, hopefully, disregarding (part of the ) calibration instructions as mentioned in the lEC

documents or not following exactly the measurement procedures as lined out in these

documents. Careful evaluation of the data and the way they were gathered seems therefore

needed in this phase of the round-robin, and will be performed in the near future.

Acknowledgements

Dr. Vobian, Deutsche Telekom -FTZ, Darmstadt, is greatfully acknowledged for the

provision of the AMOCO optical amplifier.

References

1. F.W. Willems, ''The intricacies of an EDFA round-robin: First observations from COST241 and RACE "COMFORT", in proceedings of the Symposium on Optical Fiber

Measurements, Boulder, September 1994.

2. lEC SC 86C Publication 1290-1-1, Basic specification for Optical fibre amplifiers test methods, "Test methods for gain parameters. Optical Spectrum analyzer method". Latest

version Feb. 1996.

3. lEC SC 86C Publication 1290-3-2, Basic specification for Optical amplifiers test-methods,

"Test methods for noise figure parameters. Optical Spectrum analyzer method. Latest

version, Feb .1996.

34
A GENERAL METHOD FOR MEASURING THE ELECTRIC-FIELD-DEPENDENT

ABSORPTION COEFFICIENT IN QUANTUM CONFINED STRUCTURES

IF. Siliquini, M.G. Xu and J.M. Dell

Department of Electrical and Electronic Engineering, The University of Western Australia,

Nedlands, WA, 6907, Australia

ABSTRACT

A simple method of measuring the electric-field- dependent absorption coefficient in quantum confined

structures is presented. This is achieved by measuring

the photocurrent, transmitted power and quantum

efficiency of the device and using a simple analytical

expression to extract the absorption coefficient.

INTRODUCTION

ElecLroabsorption in semiconductor multiple

quantum wells (MQWs)[l], superlattices (SLs) [2] and hetero n-i-p-i[3] structures is critical to the operation of many electro-optical. Accurate measurement of the absorption coefficient as a

function of both wavelength and applied electric field

is essential for correct design of these devices. The most direct method of measuring absorption

coefficient is by simultaneous measiu-ement of the

reflectance and transmittance[4]. While this method

is simple in concept and has been applied to MQW

devices[5], it is difficult in practice with sigmficant

error possible. Because of these difficulties the most common

technique for obtaining tlie absorption coefficient uses the measured photocurrent alone (e.g. [6,7]). This technique ignores interference effects within the

test structure which can result in large errors in the derived absorption coefficient. This technique is adequate in long AR coated devices such as waveguide modulators, but requires complete device

fabrication and careful optical alignment if results are

to be accurate. Other methods to extract absorption coefficient are based on measurement of photocurrent and

transmittance[8,9]. This technique has been applied

to MQW waveguide devices in which the incident

light is parallel to the MQW layers where

interference effects can be ignored, but again requires

careful alignment. The direct application of this technique to transverse structures results in significant errors. In this work we propose a modified method to determine the absorption coefficient by using

measured photocurrent and transmitted power in the

transverse direction as inputs into a simple closed form equation. This technique takes into account resonance effects in the test structure and requires no knowledge of the total thickness of the device after fabrication (the absorbing layer thickness is still required). An accurate method of determining the

internal quantum efficiency is also proposed. The

error analysis presented shows this method to be as accurate as the measured reflectance/transmittance

method while requiring a significantly less demanding experimental setup.

ABSORPTANCE/TRANSMISSION MODEL

A typical three layer structure used for electroabsorption measurements is shown in Fig.l.

Ohmic

contacts p-AlGaAs

and/or reflector stack n-AlGaAs

and/or reflector stack 6

Figure 1. Schematic diagram of the MQW device

used in the device model.

The middle layer is the absorbing MQW layer with

absorption coefficient a , and thickness d . The top

and bottom layers are transparent wide-bandgap

cladding layers used to form a p-i-n diode which is reversed biased to vary the electric field in the nominally imdoped middle absorbing layer. In some

structures these may be A/4 reflector stacks. To

model this device an approximation is made which

assumes that the device is an absorbing layer of

thickness d inside a Fabry-Perot cavity of arbitary finesse and free spectral range. Since the refractive index difference between the absorbing and cladding

layers is usually small, the error introduced in using

this assumption (which is verified in the error

35
analysis below) is also small. Thus, tlie 3-layered

structure (two reflector spacers plus absorbing region) can be treated as a single layer with average

refractive index, n and total thickness, d, . Under

these conditions, the normal incidence reflectance, R,

transmittance, T, and absorptance. A, can be written

as:

R =

T =

(V^ - J^e--')' + 4^RXe'^ sin' kd

(l - ^jR^e-'^J + 4^[R^e-'^ sin' kd,

(1-R,)(l-R,)e-

{l - ^R^R^e-'^))' + 4^R^R^e'^ sin' kd,

(1)

(2)

A (3) (l-R,)\l-R,e-"^-{l-R,)e-'^]

l-^R^R^e-"")' +4^R^R^e-'^ sin' kd,

where A' = 2mi/X, d is the absorbing layer thickness

and and R^ are the reflectances at the top and

bottom of the structure, respectively. The

photocurrent of the device of Fig. 1 is given by

I =n{v)^p^A (4)

p» ' he " where riiV) is the quantum efficiency (which is a

function of the applied bias), X is the signal

wavelength and Po is the incident power. The ratio of

transmitted power, P, = PT , to the photocurrent

given in Eq. (4) is tlien

P.

r](v)Ae

he

\-R,e- -1 (5)

Solving Eq. (5) for a gives

2R,

(6)

where B = hel^ / P,Ar;[v) . Equation (6) relates the

absorption coefficient to the measured photocurrent

and transmitted power. Although the calculation of a

is dependent on the absorbing layer thickness, it is independent of the total layer thickness which may

not be accurately known. Another important feature of this metliod is the fact that it is not necessary to simultaneously measure the photocurrent and

transmitted power as is tlie case in the reflectance/transmittance method of Gerber[5]. In the current model, the quantities Iph and P, can be

measured independently. Where there is a difference

in the incident power used for measuring Iph and P,, then the ratio of these incident powers need to be

included in Eq. (5).

The fundamental assumptions made in arriving at Eq. (6) are

a) Tlie refractive index is constant in llie material,

b) The quantum efficiency, t](V), is accurately

known,

c) The reflectivity, R^ , can be determined and,

d) The thickness of the absorbing layer is known.

The first assumption has two consequences: firstly, that the reflectivity does not change with wavelength,

and secondly, the only reflections of consequence are those at the air-semiconductor interfaces. The effect of this first assumption on the accuracy of the technique are a result of changes in Fabry-Perot

interference effects. An analysis of Eq. (6) shows that a is relatively insensitive to small uncertainties in R^. An assumption similar to (d) is needed in all techniques only measuring two parameters. Of the four assumptions, the accurate knowledge of rj{V) is the most difficult to determine with several workers

(erroneously) assuming this quantity to be unity [6, 7]. Determination of r]{V) is therefore very important. In the absence of significant variation in absorptance, A

the photocurrent is directly proportional to r]{U). To

a first order, the quantum efficiency can be assumed

to be wavelength independent. By choosing an

incident wavelength such that flie absorption

coefficient is known to remain high (but not necessarily constant) over the entire bias voltage range of interest, then the value of the absorptance A

will remain approximately constant. In this case, the variation of measured photocurrent with bias voltage

at constant incident power and wavelength, is a

direct measure of the quantum efficiency of the MQW photodiode.

EXPERIMENTAL EXAMPLE

To illustrate the use of fliis method, the field dependent absorption characteristics of a MQW

structure were investigated. The MQW sample used was grown by metal organic chemical vapor

deposition on a Si-doped n"" GaAs (100) substrate to the following layer specification: 0.5 jam of n'^ GaAs

(buffer), 1.0 ^m of n"" Alo.sGaojAs (etch stop), 50nm

of n Ala8Gao2As, a MQW region consisting of 50

layers of GaAs 7. Sum thick and 49 layers of

Alo.8Gao.2As 4.5nm thick, 50nm of p Alo8Gao2As,

50nm of p Alo.3Gao.7As and a cap of 20nm p^ GaAs

for contacting. Standard processing technology was

used to fabricate mesa diodes with geometry of that shown in Fig. 1. Light from a tungsten lamp was passed through a SPEX-270M monochromator, mechanically chopped

and coupled to the sample through a multimode

36
optical fibre (NA=0.1). The spectral dependence of

the power incident on tlie sample was determined by

removing the sample and measuring the photocmrent

of a calibrated Newport Ge photodiode detector in conjunction with a transimpedence pre-amplifier and

lock-in amplifier. To ensure that the power incident on the calibrated detector was the same as that incident on the test sample, the optical fibre was

placed in very close proximity (~ 0.5)a.m) to the device under test. The transmitted light was detected using the same calibrated photodiode detector as was

used to measure the incident power. The photocurrent was also measured by a transimpedence pre-amplifier

and lock-in amplifier combination. The value of

reflectivity R^was taken as that for normal incident

light at the air/semiconductor interface. The quantum efficiency was determined using the technique described above. In this case a signal

wavelength of 760nm was used where a (and

therefore -4) is expected to be high and approximately

constant with applied electric field. This was verified

after data analysis. It should be noted that this value

of A is still longer than ^direct for Alo.8Gao.2As. Figtire

2 shows the measured photocurrent, as well as, the derived quantum efficiency firom Eq. (3). As is evident from Fig. 2, the quantum efficiency

approaches unity as the bias field is increased due to carrier sweepout dominating over recombination. The

bias required to saturate the quantum efficiency to unity is approximately 6V. This is considerably

higher than the fields seen by other workers [10] (by

a factor of approximately 3) in similar MQW

structures using Alo.3Gao.7As as barriers. We attribute

this effect to the increased barrier height of the AIo gGao.aAs barriers used in oiu" case.

0 2 4 6 8 10 12 Bias voltage (V)

Figure 2 Experimental photocurrent Versus bias at X=760nm. At this wavelength the absorptance is large and constant with bias. The scale on the right

indicates the derived quantum efficiency as per Eq.

(2)

800 820 840 wavelength (nm)

Figure 3 Measured photocurrent at three values of

bias voltage.

c

90r

V—

•

u« 80 - ?

s. 70 - -0u

60 -

50 - CO

tran 40 - •0u

3 30 - CO

Met 20 -

1

1 1

v^=ov

V,-8V

/ / ...-x^ _

'r

V^=14V

1.1. 780 800 820 840 860 wavelength (nm)

Figure 4 Measured transmittedpower at three values

ofbias voltage.

This highlights the errors that are possible in assuming unity quantum efficiency in extraction of a

when using measured photocurrent, particularly at low applied bias. The measured photocurrent and

transmitted power as a function of wavelength and

applied bias, are shown in Figs. 3 and 4, respectively.

From these results the electric field dependent

absorption coefficient of tlie MQW device can be

determined. The results of using this technique are

illustrated in Fig. 5 and represents first measurement

of absorption coefficient in x=0.8 composition

GaAs/AlxGai.xAs MWQ structures. Compared with

lower GaAs/AlxGai-xAs MQWs (e.g. x=0.3) we

observe a broadening and a reduction in oscillator strength of the heavy-hole exciton peak. This may be

due to the fluctuations in the quantum well widths,

background doping and possibly tunneling into the lower X minima in the Alo8Gao.2As, which can

rapidly ionize the exciton and cause lifetime broadening[7].

37
^ 20000

c 15000

q 10000 -

5000 -

J3<

780 790 800 810 820 830 840 850 wavelength (nm)

Figure 5 Derived absorption coefficient using the measured data of Figs. (2)-(4) and using the developed model of Eq. (6). Also shown is the total error in the measured value ofabsorption coefficient.

DISCUSSION AND CONCLUSIONS

Other than errors in experimental measurements, the

principle error in determination of a is the assumption relating to tlie assumed 3 layer model of

the device. To examine tliis error an exact model that takes into account reflections of all boundaries is used

to derive the values for the transmitted power and

photocurrent. The structure considered in the exact model is identical to the MQW device measured

above. The wavelength dependent absorption

coefficient is assumed to be that shown in Fig. 5. For

the purposes of determining the extent of errors at different values of absorption coefficient, it is further assumed that the MQW layer has the same refractive index dispersion as GaAs (which is readily

available[ll,12]). The calculated values of the transmitted power and photocurrent from the exact model are then used as inputs into the single layer

model (i.e. Eq. (6)) and the estimated absorption

coefficient is calculated. This estimated value is then

compared with the initial assumed absorption

spectrum.

The error incurred in using the approximate model

together with measurement error (approximately 4%)

are shown in Fig. 5. This figure shows that the maximum error in using the model occurs when any

one of the device layers (i.e. either cladding or MQW

layer) thickness corresponds to an odd multiple of

quarter wavelength thick. At resonant wavelengths

(lie number of reflections of light at the inner

boundaries arc maximum and thus the error in using

the single layer model is maximum. Since FabiyPerot interference effects are more pronounced when

absorption is small, the errors induced by the single

layer model are largest in this region. In conclusion, a simple technique and model has

been derived which relates the measured photocurrent

and transmitted power in MQW structures to the electric-field-dependent absorption coefficient. Although the model takes into account the FabryPerot resonances in the device, it does not require

knowledge of the total device thickness nor does it require the simultaneous measurement of the photocurrent and transmitted power. This allows an

extremely simple experimental setup to be used for measurement of absorption coefficient. A simple

method of determining the electric-field-dependent

quantum efficiency using photocurrent measurements

has also been presented. Experimental results show

that the absorption coefficient can be determined

using the developed model. An initial error analysis indicates that total errors of less than 9% can be

expected when using the model where internal reflections are neglected compared to when internal

reflections are not ignored. The maximum error in using the model occurs at low values of absorption

coefficient. ACKNOWLEDGMENT : The authors wish to acknowledge the Cooperative Research Centre for Broadband Telecommunications and Networking for financial support of this project.

REFERENCES

[I] R.H. Yan, R.J. Simes, and L.A. Coldren, IEEE J. Quantum Electron., 25, 2272, 1989

[2] K.K. Law, R.H. Yan, L.A. Coldren, and J.L. Merz, Appl- Phys. Lett., 57, 1345, 1990

[3] J. Maserjian, P.O. Andersson, B.R. Hancock,

J.M. lannellis, S.T. Eng, F.J. Grunthaner, K.K. Law,

P.O. Holtz, R.J. Simes, L.A. Coldren, A.C. Gossard,

and J.L. Merz, Appl. Phys. Lett., 28, 4801, 1989

[4] Heavens, O. S., "Optical properties ofthin solid

films", Dover Publications, New York, 1991.

[5] D.S. Gerber and G.N. Maracas, IEEE J. Quantum Electron., 29, 2589, 1993

[6] P.J. Stevens, M. Whitehead, G. Parry, and K. Woodbridge, IEEE J. Quantum Electron., 24, 2007,

1988

[7] B. Pezeshki, S.M. Lord, T.B. Boykin, and J.S. Harris, Jr,Appl. Phys. Lett., 60, 2779, 1992

[8] T.H. ^ood, Appl. Phys. Lett., 48, 1413, 1986

[9] M.K. Chin, IEEE Photonic Tech. Lett., 4, 866,

1992

[10] A.M. Fox, D.A.B. Miller, G. Livescu, J.E. Cunningham and W.Y. Jan, IEEE J. Quantum

Electron., 27, 2281, 1991

[II] D.D. Sell, H.C. Casey, andK.W. Wecht.y.

Appl. Phys., 45, 2650, 1974

[12] D.E. Aspnes, S.M. Kelso, R.A. Logan, and R. Bhat, J. Appl. Phys., 60, 754, 1986

38
MEASURD4G THE RELIABILITY OF FIBER OPTIC COMPONENTS

LA. Reith

Bellcore

1. Introduction

The estimation of reliability for fiber optic components is still in its infancy. Although many failure modes

have been identified, and progress is being made in the determination of underlying failure mechanisms,

only a few studies have attempted to estimate component lifetimes or failure rates/"^ This is in part due to the immaturity of the technology. Many causes of failure have been identified with quality control or processing issues, such as poor coimector polishing quality or adhesive failure in components due to improper adhesive cure. Now that many of these failure mechanisms have been recognized and ways to avoid them identified, we are in a better position to look into long-term wear-out failures. The onset of

component degradation may not be obvious, however. Unlike active optical or electronic devices which

often show gradual performance degradation, the optical performance of passive fiber components may

not degrade at all; sudden, catastrophic failure is often the first indication that the reliability of the component is in doubt (e.g., crack-growth in fiber-based components). Due to the critical nature of

information and data travelling over today's fiber networks, prevention and/or prediction of failures is a major concern. Both laboratory and field testing are needed to provide information leading to failure prevention and to increase our confidence in the reliability of our networks.

After the component design, materials, and assembly processes have been chosen to optimize performance

and reliability, laboratory testing is the first phase of product evaluation. Most of the laboratory tests used

for fiber optic components are derived from standard tests developed for the electronics industry. While

these tests are often directly applicable to fiber optic components, there are cases where these tests are not

adequate. For example, many tests developed for adhesives do not test its compatibility with fiber or its ability to maintain dimensional stability to the sub-micron tolerances required in many fiber-optic

applications. Much work has been accomplished and is continuing in this area in standards bodies such as the Telecommunications Industries Association (TIA) and the International Electrotechnical Commission

(lEC), which develop standard fiber optic test procedures (FOTPs in the TIA) specifically for optical

fibers, components, and devices.

Regardless of how extensive and rigorous laboratory testing is, however, unanticipated field failures are hard to prevent. It is difficult to simulate actual installation procedures in the laboratory, or to predict how

different elements of the network will interact to cause failures. For example, a number of fiber breaks in mechanical splices due to torsional stresses were identified in one set of field failures. It was discovered

that the breaks were occurring due to large stresses generated when thick (900 ^m) buffer-coated fibers were looped to mount the splices into splice trays. Due to the identification and study of this problem, a new, torsional loading test was devised to better simulate actual field conditions.

In addition to installation practices, no laboratory test can replicate the combined effects of heat, humidity

bending and/or tensile stress, stress cycling, dust and dirt, condensation, and contaminants that components actually experience in the field. To this end, Bellcore has established environmental test

facilities (ETFs) in three climate extremes (hot and humid (Louisiana), bitter cold with wide temperature extremes (Maine), and hot and dry (Arizona)). These have been established for long-term testing of

components in extreme service environments. A number of passive fiber optic components, mounted in pedestals, boxes and aerial closures, are currently being optically monitored at these sites, and components

are periodically returned to Bellcore for additional measurements to ascertain degradation. In addition to optical data, temperature and humidity data are being collected.''

^

39
This paper will provide examples of some current findings in both laboratory and field testing to illustrate our approach to the determination of fiber optic component reliability. We will discuss examples including

optical cormectors, splices, and branching components.

2. Connectors

Reports of connector field failures have been rare. However, most connectors in use thus far are deployed

in protected office environments and not exposed to high humidities and large temperature swings.

Laboratory testing has shown that many connector failures are due to quality control problems related to the polishing of the connector endface. Figure 1(a) illustrates an ideal connector endface geometry where

the connector endface is radiussed with the optical fiber flush to the spherical surface. This geometry is intended to ensure that the fiber cores will always make good physical contact. Figure 1(b) illustrates a more typical endface geometry, where the spherical radius is not centered on the optical axis (apex offset), and the fiber may be protruding or recessed with respect to the spherical surface. The result of these

effects is that a gap will exist between the fiber cores. When two cormector plugs are mated, the springloading forces will cause some ferrule endface deformation and small gaps can be closed. However,

if the endface geometry parameters fall too far outside a given range, gaps will exist that can result in high

reflectance and large reflectance variation with temperature.

0.0 0.2 0.4 0.6 0.8 Fiber-Fiber Gap (|jm) (a) (b)

Figure 1 . (a) Ideal cormector endface geometry Figure 2. Reflectance as a fiinction of gap size and (b) typical endface geometry. for a typical cormector.

Figure 2 shows calculated reflectance for a connector as a fiinction of gap size at 1310 nm. Even for connectors with physical contact the reflectance is finite, due to the formation of a high-index subsurface damaged layer during the polishing process.^ As a gap opens up, the reflectance typically improves, going

through a minimum for very small gaps before beginning to increase.

Assuming that coimectors initially meet endface geometry requirements, one major reliability concern is the ability of the adhesive to maintain the fiber position relative to the ferrule within submicron tolerances

after aging in the field. Although several connectors installed at our field sites showed evidence of initial quality control problems, only one connector has shown degradation in the optical performance over time. This connector was recently removed from our field test facility in Louisiana after being on test 3 years. The connector had marginal physical contact from the outset, but the condition gradually worsened over

time. Figure 3 shows the optical reflectance performance for several days shortly after its initial installation, and over the same temperature range in the laboratory during post-mortem analyses (at the end of the 3-year period the reflectance was beyond the dynamic range of the onsite OTDR).

40
o

a:

-20 -

-40

-60

Post-Mortem Test - June. 1996

August, 1993

Time

Figure 3. Reflectance of an optical connector (a) after initial installation and (b) after 3 years in the field.

The type of behavior illustrated in Figure 3 is representative of what we have often seen in our laboratory

testing. By correlating the results of laboratory experiments with predicted gap sizes fi'om boundary

element analyses, it has been possible to determine requirements for the connector endface geometry

parameters which will ensure physical contact.^"' We have designed test fixtures in which radiussed, blank

zirconia "indentors" are mated against connectors to simulate worst-case connector loading conditions. In

this way, connector endfaces can be loaded under controlled conditions in different aging environments.

We can measure changes in the position of the fiber relative to the ferrule and measure degradation before

it shows up as an optical performance failure. We have established baseline performance data for a number of different adhesives under load and find that mechanical performance is strongly correlated to the adhesive glass transition temperature, which should be at least 20 °C above the operating temperature

of the connector for optimum performance.'" We now complete endface characterizations for all connectors newly installed at the ETFs, including measurements of ferrule radius of curvature, fiber protrusion/recess, and apex offset, using both interferometric and contacting-stylus type measurement

equipment. In future, we anticipate that comparisons of laboratory and field analyses will help us relate actual field lifetimes to accelerated laboratory testing.

3. Splices

Splices are the only type of fiber optic component deployed in the field in large numbers, and most are

fiision splices. Although mechanical splices are not as popular, they are still deployed in large numbers

compared with connectors or branching components.

Fusion splices have a very good reliability track record. Although handling operations to assemble the

splice (stripping the coating, cleaning, and cleaving) greatly weaken the fiber, the application of a typical

heat-shrink protector sleeve normally brings the splice back to a strength comparable to the pristine fiber, as illustrated by the tensile test strength distribution data plotted in Figure 6." Aging of splices in the

laboratory has shown that the integrity of the heat-shrink protector can degrade, causing the splice

strength to weaken. To date no failures have occurred at our ETFs, and, similarly to connectors, most

reported field failures have occurred due to improper assembly practices or use of inappropriate sleeve

designs.'^ New types of fusion protector sleeves of clipnon designs have recently come on the market,

however. These are typically not as strong as heat-shrink protectors, and their ultimate reliability remains

to be seen.

41
99.9

0.25 I

, ^ ^ ^ I I I I , I 10 100 1000

DYNAMIC STRENGTH, ksi

Figure 6. Fusion splice strength before (diamonds) and after (squares) application of the protector sleeve compared with the original fiber strength (circles).

There have been occasional repxjrts of field failures of mechanical splices. Most have been associated with

design flaws or improper assembly practices. For example, the series of torsional fiber breaks mentioned

in the introduction occurred when splices that did not provide strain relief were used in conjunction with

900 |im buffer-coated fiber, and the fibers were twisted as they were looped into the splice storage trays. Either a change in the way splices are mounted in the storage tray, to avoid the application of torsional

stresses, or a change in splice design, to provide strain relief, will prevent this type of failure.

Although mechanical splices come in a variety of designs, they typically are not designed to ensure good

physical contact of the fiber endfaces. Therefore, most use index-matching gels in the central aligimient

mechanism to reduce loss and reflectance, and to help protect the fibers fi'om the outside environment.

Although typical matching gels are hydrophobic, the presence of pollutants, particularly detergent or ammonia, can aid water in migrating along the glass/gel interface or diffusing into the gel.''' If mechanical splices are immersed, they can fail when water migrates around the ends of the fibers, resulting in refractive index discontinuities. There is also evidence from laboratory testing that moisture can penetrate the splice in high humidity environments and cause failures.'^ However, no performance

degradation has occurred at any of our ETFs. At present, splices in Louisiana on test for more than 3

years have had no failures.

4. Branching Devices

Like connectors, branching devices are not yet deployed in the field in large numbers. Field failures have

occurred, however, and these often have been due to fiber breakage.'^ Laboratory failures have been

reported arising primarily from materials incompatibility, often adhesives failures and mismatches in the

coefficients of thermal expansion in the packaging materials. Loss of adhesion, or bending stresses on the

fibers (sometimes resulting in fiber breakage) were typical. In one report, estimated FIT rates for 2x2

splitters were 10 to 100 times higher than the targeted FIT rates of 50.*

42
We have identified only one failure of a branching component due to fiber breakage at our ETFs, and that we believe was due to handling damage and not in-service degradation. However, we have seen some

cases in Louisiana where there have been anomalous loss variations. Figure 7 shows one example. This 3- dB coupler had increased loss during the sunmier months.

2.50-1 . 1 , \ , 1 , L

0.50 -

0 500 1000 1500 2000 Time (Arbitrary Units)

Figure 7. Loss for a 3-dB coupler on test in Louisiana from April, 1994 through February, 1996.

The branching components are test are typicaly installed back-to-back; that is the output leads are fiised in-line with each other as illustrated in Figure 8. Although the components are spaced far enough apart to measure loss and reflectance for each component individually, the OTDR measures performance across multiple legs of the component. If we assume the coupling ratio to be equally split, the actual insertion

loss can be calculated to be L = L^eas + (3/2) log2N, where L^eas is the apparent loss measured by the OTDR and N is the number of fiber branches (2 for a 3-dB coupler).^'' If the coupling ratio, excess loss, or

loss of an individual leg varies, the apparent loss will vary, but it is impossible to determine the source of

the loss variation without more in-depth diagnostic analyses more easily performed in the laboratory. At

present, we are performing such laboratory post-mortems on branching components recently removed

fi-om the Louisian ETF.

Coupler 2

Coupler 1

Figure 8. Configuration for ETF OTDR measurements.

5. Conclusions

Little data for in-service field reliability exists. To date, most field failures and many laboratory failures occur at the time of installation and can be attributed to quality control problems in either the component

or its installation. However, many of these problems have now been identified and we are beginning

longer-term studies to look for true wear-out mechanisms. This requires both laboratory and field testing.

At Bellcore we have developed several new laboratory tests, particularly for adhesives used in optical

components, that we feel are more pertinent for fiber optic component reliability than the conventional

tests used for electronics applications. In addition, we have established test facilities in three extreme

climate conditions to gain data on actual field conditions and field failures to begin to correlate laboratory

test data to in-service performance. Although much work remains, we have laid the necessary groundwork

to begin to resolve long-term reliability issues.

43
Acknowledgements

The author would like to thank E M. Vogel, O.S. Gebizlioglu, I.M. Plitz, S. Etemad, R.A. Frantz, and

H.H. Yuce for useful discussions, use of their data, and their collaborations in developing the content of

this paper.

References

' "Results of Accelerated Lifetests on Passive Optical Splitters", M. Gadonna, T. Draycott, and S. Gundersen, Proc. 13th Annual Conf. on Eur. Fibre Optic Comm. and Networks, p. 65, 1995.

- "Reliability Evaluation for PON Power Splitters", M. Gadonna, M. Redstall, and S. Gundersen, SPIE

Proc. on FO Materials and Comp., Vol. 2290, pg. 170, 1994.

^ "Reliability of Passive Fibre Optic Branching Devices", N. Fabricius, lEE Colloquim on FO Cable

Systems, No. 1994/023, 1994. 'Torsional Criteria Development for Optical Fiber Splices", S.M. Leopold, G. Kiss, A. Pellegrino, and

N. Gehlot, Proc. 7th Annual National Fiber OpticEngineers Conference (NFOEC), pg. 260, 1991.

^ "Establishment of an Environmental Test Facility for Fiber-Optic Components at Morgan City-Amelia,

LA", P. Briggs, C. Burpee, L. Slavin, J.P. Varachi, and H.H. Yuce, Proc. J 0th Annual NFOEC, pg. Ill, 1994.

^ "Construction of a Cold-Weather Environmental Test Facility for Optical Fiber, Cable, and

Components", R.A. Frantz, L.A. Reith, G. Karl, C. Weeks, D. Leavitt, and D. Barker, Proc. 11th Annual

NFOEC, pg. 734, 1995.

^ "Evaluation Method and Performance of Advanced Low-Reflection Optical Connectors", K. Kanayama,

Y. Ando, R. Nagase, and S. Iwano, Proc. 41st International Wire and Cable Symposium (IWCS), pg. 785,

1992.

^ "Highly Stable Physical-Contact Optical Fiber Connectors with Spherical Convex Ends", T. Shintaku,

E. Sugita, andR. Nagase, J. Lightwave Tech., Vol. 11, pg. 241, 1993.

^ "Effect of Ferrule-Endface Geometry on Connector Intermateability", L A. Reith, P.B. Grimado, and

J. Brickel, Proc. 1 1th Annual NFOEC, -gg. 635, 1995.

"Reliability of Epoxy Adhesives in Ceramic-Ferrule Optical Connectors", L.A. Reith, R.A. Frantz, P.B. Grimado, I.M. Plitz, O S. Gebizlioglu, and D A. Dolinoy, Proc. 43rd IWCS, pg. 790, 1994.

' "Mechanical Reliability of Fiber Optic Splices", L.A. Reith, H.H. Yuce, and P.B. Grimado, Proc. of

the Inter. Symp. Fiber Optic Networks and Video Comm., Berlin, FRG, April 5-8, pp. 316, 1993. "Analysis of Optical Interconnection Failures by Analytical Techniques", H.H. Yuce, J.P. Varachi, Jr., andT. Wei, Proc. 9th Annual NFOEC, pg. 251, 1993.

"Stability of Mechanical Splices: Void Formation and Groundwater Immersion", G.D. Kiss, E. Hershkowitz, I.M. Plitz, L.A. Reith, andE.M. Vogel, Proc. 10th Annual NFOEC, pg. 13, 1994.

"Effects of High-Humidity Temperature Cycling on Mechanical Splice Performance", L.A. Reith, R.A.

Frantz, and D A. Dolinoy, 5'P/£' />oc. 2290, pg. 139, 1994. "Passive Optical Branching Components: Increasing their Reliability", M.P Dugan, SPIE Proc. 1973,

pg. 13, 1993.

14. "OTDR Measurements Through Optical Splitters", F.P. Kapron and J.D. Berardinelli, Tech. Dig.,

Symp. Optical Fiber Measurements, pg. 7, 1992.

44
Opto-Mechanical Method for Measurement of Ferrule Concentricity and

Roundness Error

Paul Townley-Smith and Costas Saravanos

Siecor Corp., 9275 Denton Highway, Keller, TX

Abstract: A novel method for determining ferrule concentricity and roundness

error of the outside diameter is presented, based on a Fast-Fourier Transform

(FFT) of microhole position at different angular orientations of the ferrule. The

resulting concentricity value is independent of the roundness error.

I. Introduction

Connector insertion loss is one of the most important parameters in fiber-optic

transmission system design. Because concentricity error in the ferrule is a major

contributor to insertion loss, accurate measurement and control of this parameter

is very important to the connector manufacturing industry. Currently, all measurements of concentricity error are influenced by the roundness error of the

outside diameter of the ferrule, which reduces the precision and accuracy. At the

same time ferrule roundness error is an important parameter that characterizes

the quality of the ferrule. A separation between ferrule concentricity and

roundness error will improve the control of connector insertion loss and identify

ferrules with poor roundness error. A standard opto-mechanical method for

ferrule concentricity measurementt^'^l positions a ferrule in a v-groove as shown

in Figure 1. The position of the center of the microhole is measured, then the

part is rotated through 180° and the position of the microhole is measured again.

The difference between the two microhole positions is calculated as concentricity.

This measurement is accurate and precise when the roundness error of the

outside diameter is small compared to the concentricity. As the roundness error increases the concentricity measurement accuracy reduces because the roundness

error contributes to the movement of the ferrule microhole as it is rotated. Other

standard methodsl^l improve the measurements precision by monitoring the

position of the center of the microhole at several different angular orientations of

the ferrule. However the accuracy would still suffer because of the contribution

of roundness error. This paper proposes that an appropriate application of a FFT

to the microhole position data can be used to separate pure concentricity and

roundness error achieving improved precision and accuracy in the measurement

of concentricity while obtaining an accurate roundness error measurement.

II. Theory

For a perfectly round ferrule rotating in a v-groove as shown in figure 2, the

center of the microhole will trace out a path given by

45
^ = ^o+[^Jsin^ (1)

y = yo+[^oos0 (2)

where and are the coordinates of the center of the outside diameter of the

ferrule, c is the concentricity of the ferrule, 0 is the current rotation angle of the

ferrule. For a ferrule that has an outside diameter that is out of round, we can

define the radius r{0) as

ri0) = r,+S(O) (3)

where is the mean radius of the ferrule and S{0) is the roundness error. As

long as 1^(^)1 any roundness error will manifest itself as an offset

perpendicular to the surface of the v-groove. For a 90° v-groove, the x and y

offsets due to roundness error will be orthogonal to each other, and will also be

related by a 7r/2 phase factor. Thus in the presence of roundness error, the

center of the microhole will trace out a path given by

x(O) = x,+[^smd + S{0) (4)

y(0) = yo+[^cosd + s{^0 + ^

as the ferrule is rotated in the v-groove. N consecutive discrete values of x and y

coordinates of the center of the microhole are taken equally spaced in 0 and

covering one complete revolution of the part. The FFT of the x or y data will

have a magnitude of c/2 at the one cycle per revolution frequency component,

which is interpreted as concentricity. The Inverse FFT (IFFT) of the transformed

data with the concentricity and mean offset components removed, will yield the

roundness error traces. From these traces, the peak-to-valley error, lobe

frequencies or other useful roundness error measures can be obtained. The

concentricity and roundness error estimates based on x and y data can be

averaged together to improve measurement precision.

III. Experimental

All measurements were made with an apparatus as shown in Figure 1. The

ferrule is backlit with a high power white light source, and rotated in the v- groove. A video camera is attached to a high power microscope. The video

images are processed in the computer to calculate the center of the microhole.

To examine the accuracy of the roundness error measurement, several ferrules

were measured using the proposed method and using a stylus type mechanical

roundness error measurement system. The comparison of the measurements is shown in Figure 3, and demonstrates the high correlation between these two

instruments. Figure 4 compares the raw data trace, the FFT roundness error and

46
the stylus-type roundness error traces for one part. As seen in the figure, the raw

trace does not directly reveal the concentricity and roundness error of the part.

However, an accurate roundness error measurement is obtained using the FFT

method.

Figure 5 is a demonstration of the 2-point methods accuracy problems. A few

ferrules with varying roundness error to concentricity ratio were measured. The

ferrules were measured several times using the 2-point method and once using

the FFT method. The average of all measurements using the 2-point method

were recorded and the difference from FFT-based concentricity was calculated.

As shown in the figure, the two methods converge for low roundness

error/ concentricity ratios as expected, but diverge rapidly at ratios higher than

0.5.

rv. Conclusions

A new method for measuring ceramic ferrule roundness error and concentricity

has been presented. This method can separate the concentricity and roundness

error of ferrules using an FFT algorithm. In that way, the measurement accuracy,

repeatability and correlation to connector insertion loss is improved significantly.

V. References

1. lEC Publication 1300-3-25 "Concentricity of Ferrules", International

Electrotechnical Commission, 3, rue de Varembe, Geneva, Switzerland

2. Pasturczyk, Z.; Wong, B.; Saravanos, C, "Fiber Optic Geometry Test Station"

Proceedings of the IWCS, pp. 623-626, 1989

3. TIA/ElA-455-135, "Connector Ferrule Inside and Outside Diameter Circular

Runout," Telecommunications Industry Association-Electronic Industries

Association, 2500 Wilson Blvd., Suite 300, Arlington VA 22201.

CCD

ligjTt Source FerruleMder NScrasoope OMnputer

Figure 1. Ferrule Measurement Apparatus. Figure 2. Ferrule in V-groove

47
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7

Stylus P-V Roundness Error (um)

Figure 3. FFT roundness error vs stylus type roundness error

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 Ferrule Roundness Error/Cone Ratio

Figure 5. Difference between two point and FFT Concentricity

48
Errors due to connectors in optical fiber power meters

Igor Vayshenker, Xiaoyu Li, Darryl A. Keenan, and Thomas R. Scott

National Institute of Standards and Technology, 325 Broadway, Boulder, CO 80303

Abstract

We discuss results of a major potential error source in the use of optical fiber power meters:

outputs can vary dramatically when using various types of connectors or even connectors of the

same type but from different vendors. We investigate the magnitude of this connector-induced

variation by calibrating several types of optical fiber power meters at three telecommunications

wavelengths of 850, 1310, and 1550 nm. In these measurements we vary the connector type and

connector vendor, and observe the resulting offsets in calibration resuhs. Observed variations of as much as 1 0% were found, due, presumably, to the different reflection properties of the detectors,

windows, and connectors involved. A test meter user, therefore, can expect an error as large as 10%

if the optical fiber power meter is used with a different connector (or vendor) than that used for

calibration.

1. Introduction

When optical fiber power is measured, radiation is transmitted to an optical fiber power

meter through a fiber attached to a detector by a fiber connector and adapter. The proximity of a

fiber connector to a detector and its associated window provide an opportunity for reflections to introduce errors in the power readings.''^ Even though the measurements using connectors are generally repeatable, the connectors can skew the measurements results. To investigate this issue

we conducted a study involving the calibration of optical fiber power meters using various types of

connectors. We selected six common connector types: FC/PC, FC/APC, ST, biconic, SC, and SMA

from four vendors chosen randomly (the vendors are identified by letters A through D). Calibrations were performed on four types of power meters (the meters are identified by numbers 1 through 4)

at three telecommunications wavelengths: 850, 1310, and 1550 nm.

2. Measurement system

The system depicted in Figure 1 was used to perform the calibrations. As a reference we used

a commercially available, electrically calibrated pyroelectric radiometer (ECPR).^ The detector is approximately 8 mm in diameter and has very low reflectance over the wavelength range. The

pyroelectric sensor is made of lithium tantalate and is covered with a gold-black coating.

The calibration system contains three laser sources and a positioning stage for comparing

the outputs of the ECPR and the test meter. All the optical fibers in the system are single-mode.

Each laser source plate contains a laser diode whose output is transmitted through a fiber to a fiber

splitter from which about 1% of the energy travels through a fiber to a monitor. Power changes (for

example, due to diode laser instability) are taken into account by the monitor readings. The fiber

jumper cables are stationary during the calibrations. Power is coupled directly into the pyroelectric

detector or the test meter. The incident power was about 100 |iW.

49
Digital

Voltmeter

Temperature

Coitroller

Laser Power Cuont

Sipply

Rt^le

Mountiiig PtetfOTm

ISSOmn

Plate NISTECPR

ion j System

User's fiber and comKctor

BCD/GPU

Coivegter

NIST Standard (ECPR)

Uscflest Meter

Figure 1 . The measurement system.

Table I. Connector study variables.

Connector Types: FC/PC, FC/APC, ST, biconic, SC, and SMA

Connector Vendors: A-D

Wavelength: 853, 1307, and 1549 nm

Meter/Detector Type: Meter 1 (Si & Ge remote sensors; angled), Meter 2 (InGaAs fiber-pigtailed

sensor), Meter 3 (Ge sensor with window; not angled), and Meter 4 (windowless Ge sensor; not

angled).

3. Results

The measurement variables are given in Table I; the results are reported in the form of

Figures 1-8. Due to the format of this paper, we included only a few of the measurement results. We performed 5 data runs for each measurement; the data show that the repeatability of

measurements performed with a specific connector is reliable. The connector's offset is consistent.

Figures 1-4 depict the results for four power meters and six connectors referenced to an FC/PC

connector for vendor C. The largest differences occurred for the germanium power meter with a window (the sensor was not angled). Placing a sensor at a small angle will decrease the reflection

between the sensor, the sensor window, and the connector. The observed differences were more

than 10%. Figures 5-8 depict the results for four power meters and four connector vendors

50
referenced to ST connector for vendor A. Again, the largest differences (8%) were observed for

the germanium power meter with a window. Vendor B's connectors caused a greater offset than that

caused by connectors of all other vendors.

Wavelength (nm) Wavelength (nm)

-IS- FC/APC ST -—-Eicon _^FC/APC-^ST --- Eicon

_^SC ^SMA -^SC -^SMA

Figure 2. Si & Ge power meter (angled); referenced Figure 3. InGaAs power meter (flber-pigtailed);

to FC/PC for vendor C. referenced to FC/PC for vendor C.

8

6

4

& 2

-2-4-6

1 307 1 549

Wavelength (nm)

FC/APC ST -m- BIcon SC -e- SMA

Figure 4. Ge power meter with window (not

angled); referenced to FC/PC for vendor C.

1307 1549

Wavelength (nm)

FC/APC ST BIcon SC SMA

Figure 5. Ge power meter without window (not

angled); referenced to FC/PC for vendor C.

0.2

0.1

o.o

-0.1

-0.2 853 1307

Wavelength (nm)

• Vendor B Vendor C

1549

• Vendor D

Figure 6. Si & Ge power meter (angled);

referenced to ST for vendor A.

853 1307

Wavelength (nm)

Vendor B Vendor C

1549

• Vendor D

Figure 7. InGaAs power meter (fiber-pigtailed);

referenced to ST for vendor A.

51
-2 1307 1549

Wavelength (nm)

--Vendor B -^Vendor C -^Vendor D Vendor B Vendor C Vendor D

Figure 8. Ge power meter with window (not

angled); referenced to ST for vendor A.

Figure 9. Ge power meter without window (not

angled); referenced to ST for vendor A.

4. Conclusions

We found significant measurement offsets resulting from the use of various connectors and

a variability within a single connector type obtained from different vendors. Thus, errors could

likely occur when changing types of connectors or connector vendors on fibers connected to optical

fiber power meters. For accurate measurements during the calibration of the instrument, we suggest

that the meter owner should be aware of the connector's effects. A transfer standard, whose output

is insensitive to the connector types, is a usefiil tool when determining effects due to various

connectors. The connector adapter used with the connector is an integral part of the measurement.

The magnitude of the errors is wavelength-dependent. The offset is small if a connector does

not have a reflecting surface or a power meter sensor is angled. A connector with a reflecting

surface will cause the meter to read incorrectly, usually higher. It is very important to calibrate an

optical fiber power meter with the same type of a connector used in the actual measurement.

This work is supported by the Calibration Coordination Group (CCG) of the Department of

Defense; the lead agency for this project is the U.S. Army, Huntsville, Alabama.

1. R.L. Gallawa and Xiaoyu Li, "Calibration of optical fiber power meters: the effect of

connectors," Applied Optics, Vol. 26, No. 7, pp. 1170-1174, April 1, 1987.

2. Xiaoyu Li and R.L. Gallawa, "Calibrated optical fiber power meters: errors due to variations

in connectors," Fiber and Integrated Optics, Vol. 7, pp. 241-248, March 3, 1988.

3. R.J. Phelan, Jr. and A.R. Cook, "Electrically Calibrated Pyroelectric Optical-Radiation

Detector," Applied Optics, Vol. 12, p. 2494, October 1973.

Acknowledgments

References

52
MU-type PANDA Fiber Connector

R. Nagase and S. Mitachi

NTT Opto-electronics Laboratories

Tokai, Naka-gun, Ibaraki 319-11, Japan

Introduction

Polarization-maintaining fibers (PMFs) e. g., PANDA fiber[l], are widely used in the field of

optical sensing systems[2] and advanced telecommunication systems[3]. We require high

performance and compact PMF connectors to improve these practical systems. However, there is no PMF connector smaller than the SC connector[4] because of the design difficulty involved in precisely matching the principal axes of connected PMFs.

In this paper we propose a keying accuracy requirement for PMF connectors which will allow the

realization of PMF connection with high extinction ratio propagation. And also we describe the

design principle and performance of a compact, newly developed MU-type PANDA fiber connector.

Definitions and measurement of the keying accuracy and the extinction ratio

Figure 1 shows an experimental configuration for measuring the direction of the principal axis

(PA) of a PMF and its extinction ratio. Linearly polarized light with a constant power in any

direction can be produced by using a setup consisting of a polarizer (Pi ), a quarter wave plate (Q)

and a polarizer (P2). The patch cord under test is measured in the following way. First, an analyzer

(A) and a polarizer (P2) are adjusted so that the transmitted light power has a minimum value of

Pmin (dBm). The PAs of connectors Ci and C2 are parallel to the principal axes of P2 and A,

respectively. Then analyzer A is adjusted so that the transmitted power has a maximum value of

Pmax (dBm). We define the extinction ratio (ER) of the output signal as Pmax - Pmin in dB units.

The value of ER changes along the fiber because of the modal birefringence.

Keying accuracy requirement

The minimum extinction ratio £hnin for a PA misalignment angle 0 between connected PMFs is given by[5]

£rmin = lOlog(tan^0). (1)

To achieve a minimum ER of 20 dB, the allowable misalignment angle is estimated to be 6° from

eq. (1), i. e., the keying accuracy should be ±3° (Fig. 3). If two PMF patch cords are connected,

the minimum total extinction ratio Ertmm will be

Ertmxn = -201og(10-^^"'° +

^Q-Enno ^ ^^^^^ (2)

where £ri and En are the individual ER values of two patch cords. A typical patch cord with a several-meter-long PMF and two connectors has an ER of 25 ~ 35 dB.

Figure 4 shows the measured ER of randomly concatenated PMF connectors whose keying

accuracies were better than ±3°. We confirmed that all measured ER values were higher than the

calculated Ertmin from eq. (2), and a keying accuracy of ±3° is sufficient to achieve an ER of 20

dB.

53
MU-type PANDA fiber connector design

( 1 ) Keying accuracy: PANDA fiber has a clear geometric axis (GA) parallel to the stress applying

parts (SAPs) as shown in Fig. 2. It is known that the angle between the PA and the GA is less than

r[6]. It is much easier to observe the GAs by microscope than to measure the PA directions.

Therefore, we have established a keying accuracy of ±2° for GAs.

(2) Keying direction: A PANDA fiber has an elliptic cross section whose major axis is parallel to the y-axis (as shown in Fig. 2) because of the SAPs shrink as the fiber is drawn. Therefore, after

it has been glued to the ferrule, the fiber core has a tendency to shift in the x-axis direction. Figure

5 shows the measured fiber core eccentricity directions. The key direction (0°) was set in the x-axis

direction. The distribution will usually be uniform, however, in this case about 90% of the core

direction distribution was localized in the ±60° region. It is desirable to achieve a low connection

loss[7], therefore, the keying direction should be set in the x-axis direction as shown in Fig. 2.

(3) Ferrule structure: A ferrule floating mechanism is essential for realizing stable connection. In

particular, the floating distance cannot be reduced if the MU connector is used as a backplane

connector[8]. Therefore, we have designed a new MU-type ferrule with a specially designed

coupling device as shown in Fig. 6. The ferrule can float from the plug housing by the same

distance as a normal MU plug, but it cannot rotate around the ferrule axis. The flange can be

pressed into the ferrule in any direction after the fiber has been glued to the ferrule.

(4) Plug housing: The rectangular shaped design and the push-pull latching mechanism used for

MU connectors are suitable for PMF connectors because no torque is applied to the ferrule during

the coupling operation.

Characteristics of MU-type PANDA fiber connector

We assembled 2 m-long MU-SC patch cords with 10 (im MFD PANDA fibers and measured them

at an 1310 nm LD light source. Figure 7 shows an MU plug attached to a PANDA fiber. We confirmed that the keying accuracy for the GA of a PANDA fiber can be adjusted to within ±2°

of plug key using microscope observations of the SAPs and flange pressing technique as shown in

Fig. 8 (a). The average ER was 27.3 dB for random concatenations of 2 patch cords as shown in Fig. 8 (b). Figure 8 (c) shows the connection repeatability. The ER fluctuations were less than 8

dB for 500 mate-demate cycles. The average connection loss was 0.11 dB for random

concatenations as shown in Fig. 8 (d). The return loss of a connection point was typically 50 dB.

These values are similar to those of normal MU connectors for SM fibers.

Conclusion

We investigated the extinction ratio degradation of PMF connections and determined a keying

accuracy of ±2° for geometric axes and the desirable keying direction for PANDA fiber connectors.

We have also developed a compact MU-type PANDA fiber connector which can be used for both

fiber cable coupling and as a backplane connector. It features a high extinction ratio for random

connections (typically over 20 dB), good connection repeatability, and a low connection loss

(typicallyO.il dB).

References

[1] T Hosaka, et al.. Electron. Lett., 21, no. 20, pp. 920-921, 1985.

[2] R. Ulrich, etal.. Opt. Lett., \bl. 4, pp. 152-154, 1979.

[3] T. Kataoka, et al., Electron. Lett., \bl. 27, no. 1 1, pp. 943-944, 1991.

[4] R. Nagase, et al., in Proc. lEICE General Conf., Japan, paper C-280, 1991 (in Japanese).

54
[5] I. Yokohama, et al., IEEE J. Lightwave TechnoL, Vol. LT-5, no. 7, pp. 910-915, 1987.

[6] Y. Kikuchi,et al., in Proc. lOOC '89, paper 19B3-12, 1989.

[7] Y. Ando, et al., lEICE Trans., Vol. E77-C, No. 12, pp. 1970-1982, 1994.

[8] S. Iwano, et a.\.,IEEEJ. Lightwave Tech., vol. 10, no. 10, pp. 1356-1362, 1992.

0^

test S Li Pi Q P2 1^ Ci patchtchcord cor C2 L3 A D

Fig. 1 . The experinnental configuration for the direction of PA and the ER measurement

stress applying

part (SAP)

ferrule plug PANDA fiber

Fig. 2. PANDA fiber and its connector plug geometry

principal y geometric axes \iflaxes

keying

accuracy

Fig. 3. Definition of keying accuracy

15 20 25 30

calculated minimum extinction ratio (dB)

o

c

0)

CD

40

30

20

10

0

key

ferrule center

core rn 1^i

n =236

ave.= -0.2° S.D.= 34.7° n /

/ /

/ M

-90 -70 -50 -30 -10 10 30 50 70 90°

-80 -60 -40 -20 0 20 40 60 80

fiber core direction (deg.)

Fig. 4. Experimental results for randomly

concatenated PMF connectors

Fig. 5. Distribution of the fiber core

eccentricity directions

zirconia tip coupling device flange

5

coupling device floating direction from plug housing

ferrule floating direction from coupling device 0ferrule floating from plug housing

Fig. 6. Structure of the ferrule for an MU-type PANDA fiber connector

55
Fig. 7. MU-type PANDA fiber connector plug

7-7

i

n = 20

ave. - -0.4° min = -1.4° max = 0.8°

-5 -4 -3 -2 -1 0 1 2 3 4 5

keying accuracy for geometric axis (deg.)

(a)

70

60 ^

50

§ 40

S" 30

20

10

0

n = 380

ave. = 27.3 dB

S.D. = 4.33 dB

12 16 20 24 28 32 36 40 44

extinction ratio (dB)

(b)

200 300

mating cycles

(c)

500

n = 190

ave. = 0.11 dB

S.D. = 0.10 dB

—^1—r"i F"

0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 connection loss (dB)

(d)

Fig. 8. Characteristics of MU-type PANDA fiber connectors; (a) keying accuracy of the PANDA

fiber geometric axes, (b) measured extinction ratio of randomly concatenated connector pairs,

(c) connection repeatability, and (d) connection loss of randomly concatenated connector pairs.

56
High Performance Single -Mode wavelength independent

all-fiber build-out attenuator

Shin-Lo Chia

AMP Passive Products Division

2525 E. Bayshore Rd. Palo Alto, CA 94303

Tel : 415-493-8008 FAX: 415-493-8924

e-maU: AMPMS.kap00110@AMPUSEMl.amp.com

Abstract

A very high performance, cost-effective, miniature, wavelength independent all-fiber Single

Mode attenuator is demonstrated. The device maintains consistent attenuation in both jumper-tojumper, and direct-to-detector applications.

Summary

In this paper a novel, all-fiber attenuator that uses a high attenuation core single-mode fiber is proposed and demonstrated. The basic process is as follows: (1) A length of standard singlemode fiber is fiision spliced to the attenuation fiber. (2) The attenuation fiber is cleaved at a precisely controlled length which will yield the required attenuation. (3) Another fiber is fiised to the other end of the attenuation fiber [ See Figure 1 (a) ]. (4) The finished attenuator is assembled

into the ferrules and PC polished [ See Figure 1 (b) ]. (5) The ferrule assembly is packaged as a build-out attenuator [ See Figure 1 (c) ].

The net attenuation provided by the proposed design depends on the following factors:

1 . Absorption by the dopant material in the core.

2. Overlap integral between the mode field and the physical attenuation core.

3. Mode field mismatch at each fiasion junction.

4. Modal interference and noise generated by the four optical interfaces in the attenuator

system.

In order to determine the specific characteristics required of the core dopant, a beam propagation

model was constructed. In the model simulation, from 0.0 to 5.0 mm is the standard S.M. fiber. From 5.0 to 15.0 mm distance, the absorption coefficient of the core is assumed to be wavelength

independent, and is adjusted to produce an attenuation of 1 dB/mm with a 6 \im diameter core. Then from 15.0 to 20.0 mm the standard S.M. fiber is replicated. As expected, the model exhibits a pronounced wavelength dependence effect, showing a decreasing loss vs. wavelength. (Refer to

Fig. 2) This arises from the fact that the longer wavelengths with larger mode field diameters will

57
decrease the overlap integral (attenuation) between the mode field and the attenuation core. In

order to achieve true wavelength independence in the finished device, this effect must be

compensated for by the wavelength response characteristics of the proposed attenuating fiber.

Figure 3 shows the total contribution effect of the wavelength dependency for the attenuation

fiber BOA. The center line shows the normalized measurment BOA spectral response. A

attenuation value increase with the increase of wavelength . This response is balanced between the

dopant absorption and the mode propagation loss dependence. A third modal noise ripple effect

could add to the balanced response and make the final BOA spectral with the ripple on. However, the carefully design the attenuation fiber with better matching characteristics and well

control the fusion quality between the attenuation fiber and standard fiber could dramatically

decrease the modal noise effect.

The key component in this attenuator is a specialized attenuation fiber which has been

manufactured to meet these specific criteria. The attenuation fiber is manufactured by using a GeAl co-doped silicate core and introducing transition metals [4] or rare earth elements to the core. A difficulty arises when increasing the dopant concentration in order to achieve high attenuation

per unit length as required when fabricating a very short length attenuator. High dopant

concentration may increase the NA value dramatically and also reduce the modal noise effect; but

increase backreflection and cut-off wavelength as well. Current research centers on the

determination of the optimal characteristics required of the attenuation fiber. When these

parameters are finalized, and fiber is produced to these specifications, manufacturing will become

a simple matter of controlling the quality of the fusion splice, and the accuracy of the length of the

attenuation fiber. These process readily lend themselves to cost-effective mass production.

An additional undesirable side effect of the mismatch in NA between the two fiber types is a discrepancy in effective attenuation between an application where the attenuator is used in line

between two single-mode cables and when the output of the attenuator is coupled directly into a

detector. In the latter case, any light which has scattered into the cladding can represent a

significant portion of the output signal and reduce the originally designed attenuation level

especially in high attenuation level ( 15 dB and above ). A significant reduction of these cladding

modes can be effected by chemically etching away some of the cladding of the attenuation fiber.

After the first optimized attenuation fiber was obtained and tested, an entire range of attenuators

was successfully fabricated and tested. The current precision fusion and cleaving process can be

controlled by ± 50 |am, allowing accurate production of attenuators with values as low as 0.75 dB

or as much as 25 dB simply by controlling the length of the attenuating fiber included in the

device. Figure 4 shows a typical wavelength scan for such an attenuator. It can be seen that all meet attenuator tolerance requirements within the 1260 to 1360 nm and 1430 to 1580 nm

passbands as specified in Bellcore TA-TSY-000910.

When fusing the standard fiber to the attenuation fiber, the fusion joint has no more than -65 dB

backreflection. Therefore, the backreflection of the build-out attenuator as shown in Figure 1 is not limited by the attenuator element itself but by the polishing quality of the ferrule ends.

58
Currently, a new Ultra-PC polishing process better than 55 dB in return loss and with less than 50

nm fiber undercut [6] has been successfully demonstrated and implemented in this BOA product.

Conclusions

A very ragged and high performance all fiber attenuator that revolutionize installation and

operation simplicity has been obtained. The manufacturing is just a simple matter of controlling

the quality of the fiasion splice, and the accuracy of the length of the attenuation fiber. These

processes readily lend themselves to cost-effective mass production. Also, the new attenuator

meets all Bellcore requirements for attenuator tolerance, backreflection, polarization stability,

humidity resistance, etc. References

1. J.C. Goodwin and P.J. Vella, " Modal noise in short fiber section "

,J. Lightwave Tech.,

Vol. 9, pp. 954-958, 1991.

2. S. Heckma, " Modal noise in single mode operated slightly above cut-off Electron letter. Vol. 17, pp. 499-500, 1981

3. R. A. Throckorton, D. Harris, " Modal interference in field installable single-mode fiberoptic connector" NFOEC 94, Vol.3, pp. 399-406, San Diego, 1994

4. P. C. Schultz, " Optical absorpfion of the transition elements in vitreous silica ", J. of the

American ceramic society. Vol. 57, No 7, pp. 309-3 13, 1974

5. "Generic Requirements for Fiber Opfic Attenuators", TR-TSY-000910 Issue 2, Sep. 1992.

6. "Generic Requirements for Single-Mode Optical Fiber Connector", GR-326-CORE Issue 1, Dec. 1994.

59
BPMCAD Simulation

Attenuated Fiber ( 1 dB/mm Att.)

iso-boabpm4

0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20

Propagation Distance ( mm ) — 1250.00 —e— 1300.00 1350.00 --^ 1400.00 —^ 1450.00 1500.00 1550.00

Figure 2. Graph shows the calculated results of the BPM model simulation. From 0.0 to 5 .0 mm is the standard S.M. fiber. From 5 .0 to 15 .0 mm distance, the absorption coefficient of the core is assumed to be wavelength independent, and is adjusted to produce an attenuation of 1 dB/mm a 6 |im diameter core. Then from 15.0 to 20.0 mm the standard S.M. fiber is replicated.

Wavelength ( nm)

Figure 3. Total wavelength dependency effects for the attenuation fiber BOA

60
Nonlinear coefficient of optical fibers at 1550 nm

Valeria L. da Silva, Y. Liu, AJ. Antos, G.E. Berkey and M.A. Newhouse

Corning Incorporated, Science and Technology Division,

Sullivan Parl<, NY 14831

A major revolution in lightwave transmission system is ocurring with the advent of

erbium-doped fiber amplifiers. They have allowed transmission over transoceanic distances

without the need of electronic repeaters and enabled high speed, multiwavelength

terrestrial transmission systems. On the other hand, the use of optical ampUfiers has

increased the optical signal power to a point where nonlinear effects start to become

important and impose serious limitations to system performance. The magnitude of the

major nonlinear effects, self-phase modulation (SPM) and four-wave mixing (FWM),

depend on the nonlinear coefficient n2/Aeff (nonUnear refractive index/effective area).

Therefore it is important to measure the n2/A€ff for telecommunication fibers. In this paper

we wiU review measurements of the nonlinear coefficient by different laboratories and

present results obtained at our laboratory on different types of fibers of interest in

transmission systems at 1550 nm.

The value of n2/Acff is usually obtained by measuring the magnitude of nonlinear

effects such as self-phase modulation, cross-phase modulation or four-wave mixing. By an

independent measurement of the Aes, n2 can be obtained. Almost all measurements of n2

for non-polarization-preserving dispersion-shifted fibers resulted in a value of 2.3 ± 0.1 x

61
10'^° m^AV at 1.5 |im [1-6]. The exceptions were cross-phase modulation measurements in

which the pump source is slowly modulated. In these cases, higher n2 values were

obtained, 2.7 - 2.9 x

10"^° m^AV [7,8]. The higher n2 is attributed to the electrostriction

contribution, which becomes significant for pulsewidths of ~1 ns or longer [9].

Current lightwave systems use other fibers besides dispersion-shifted fibers. Since

n2 depends on the amount of GeOa and the mode-field profile, it is important to measure

Til in the different fibers. We measured the n2/Acff coefficient, using the self-phase

modulation technique [2,11], for three different fibers: dispersion-shifted fiber (DSF),

large effective area fiber (LEAF) [10] and dispersion compensating fiber (DSF). The

characteristics of the measured fibers are summarized in Table 1.

Fiber

Type

Length

(km)

'^meas

(nm)

a @ 1550 nm

(dB/km) (nm)

Dispersion

(ps/km/nm)

MFD

(^im)

Aeff

DSF 5.0 1545 0.21 1547 -0.16 8.32 51

DCFl 1.0 1550 0.50 -88.7 5.98 29

DCF2 1.0 1550 0.85 -125.6 5.67 28

LEAF 2.9 1535 0.21 1538 -0.37 9.79 88

Table 1. Characteristics of tlie measured optical flbers. The effective areas were measured by the

far-field pattern method.

The major effect of SPM is to broaden the spectra of input pulses at high input

powers. The amount of broadening depends on the fiber length, peak input power and

fiber dispersion. In our measurements, we used relatively long pulses (-40 ps) and short

fiber lengths in order to avoid any dispersion effects. Under these conditions, the spectral

broadening factor for a Gaussian pulse is given by [1 1]:

62
(AO)),

1 +

3V3

271: PL (1)

where Aco is the RMS spectral width, \ is the wavelength, P is the peak input power and

Leff is the effective length.

We measured the output spectrum as a function of the peak input power utilizing

the experimental setup depicted in Fig. 1 . The pulse source consisted of a mode-locked

tunable external cavity laser producing 40 ps pulses at 5.0 GHz repetition rate followed by

an erbium-doped fiber power amplifier with saturated output power of 20 dBm. In order

to increase the peak power od the pulses, an electro-optic modulator with a duty cycle of

-10% was introduced prior to the optical fiber amplifier. The resulting pulses had a peak

power in excess of 2.5 Watts allowing the use of short fiber lengths (1-5 km) and reducing

the effect of fiber dispersion.

ML-LASER MODULATOR -

Synthesizer

TAP

Test Fiber

TAP

Power

Meter

2:1

SCANNING

FABRY-PEROT

Figure 1. Experimental setup.

The spectral broadening factor (Ao))out/(Afi))in as a function of the peak input power is

shown in Fig. 2 for the measured optical fibers; the ratio between the DSF value and the

other fibers is shown in Table 2. As expected, the spectral broadening is less severe in

63
LEAP than in DSF. Our results show that n2/Acff of LEAF is ~L7 times smaller than

DSF wich is in very good agreement with the ratio between measured effective areas.

5 I—'—I—"—I—I—I——I—I—I—I-—!—I—I—•—I—

O

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5

Input Power X Effective Length (W x km)

Figure 2. Ratio of output and input RMS spectral widths as a function of peak input power times

tlie effective length for a dispersion-shifted fiber (DSF), large effective area dispersion-shifted fiber

(LEAF) and dispersion compensating fibers (DCF). Dashed lines are a fitting to £q. 1.

Fiber Type Aeff/(Aeff)^'^ (n2/Aeff)^'^/(n2/Aeff)

DCFl 0.56 0.53

DCF2 0.54 0.53

LEAF 1.72 L70

Table 2. Measured values of Aeir and tiz/A^a normalized to DSF values. The n2/Acff of DCFs is expected to be higher than in DSFs because of the smaller

Aeff and higher GeOi content. It was measured to be -3.0 times higher than in DSF in Ref.

[71 and between 2.4 - 3.5 times higher in Ref. [61. Our measurements of two different

64
DCFs, however, show an increase of only 1.9 times in the n2/Aeff, which agrees very well

with the decrease of -1.8 in the Aeff . This indicates that, unlike the DCFs used in Ref. [6]

and [7], our DCFs have an n2 comparable to DSFs.

REFERENCES:

[I] L. Pringent and J.P. Hamaide, IEEE Photon. Technol. Lett., vol. 5, pp. 1092

(1993) .

[2] R.H. Stolen, K.S. Kim, S.E. Evangelides, G.T. Harvey, W.A. Reed, OFC'95,

Paper FD-1, San Diego, CA.

[3] Y. Namihira, A. Miyata and N. Tanahashi, Electron. Lett. vol. 30, pp. 1 171

(1994).

[4] M. Artiglia, E. Ciaramella and B. Sordo, Electron. Lett. vol. 31, p. 1012 (1995).

[5] S. Chemikov and J.R. Taylor, OFC'96, paper ThS5.

[6] A. Boskovic et al., CLEO'96, paper CMA7.

[7] T. Kato, Y. Suetsugu, M. Takagi, E. Sasaoka, and M. Nishimura, Opt. Lett.

vol.20, pp. 988 (1995).

[8] A. Fellegara et al., OFC'95, paper ThS4.

[9] E.L. Buckland and R.W. Boyd, OFC'96, paper ThF2.

[10] Y. Liu, V.L. da Silva, A.J. Antos, V.A. Bhagavatula, D.Q. Chowdhury, and M.A.

Newhouse, IOOC'95, Paper PD-19, Hong Kong.

[II] G.P. Agrawal, "Nonlinear fiber optics". Academic Press, 1989,

65

Polarization independent nonlinear refractive index

measurement in optical fiber

Andrea Melloni and Mario Martinelli

Politecnico di Milano, Dip. Elettronica e Informazione

Andrea Del Core and Andrea Fellegara

ITALTEL, BUTE, Castelletto di Settimo Milanese

Abstract

Two techniques for the polarization independent measurement of the nonlinear

refractive index n2 of optical fibers for telecommunications are presented. They allow

accurate measurements of 712 and the evaluation of the nonlinear response of any

optical fiber employed in transmission optical systems.

Introduction

Optical fiber nonlinearities can represent a serious drawback in optically amplified transmission systems. The availability of a simple and accurate method for the measurement of

the nonlinear refractive index (or n2/Aeff) is a really important issue, especially since

when optical fiber providers offer a large varieties of special fibers with particular nonlinear

characteristics, such as doped fiber or fibers with very large or very small effective area. The value of n2 can be recovered by various methods: Self-Phase Modulation (SPM) [1], Cross-Phase Modulation (XPM) [2, 3] and Modulation Instability (MI) [4]. Methods based on SPM determine the distortion of a very short optical pulse after the propagation in the

fiber and those based on XPM measure the phase shift induced on a weak probe beam

by an intense optical beam (the pump). Unfortunately all these methods are polarization

dependent and the recovered value of n2 contains an uncertainty due to the State of Polarization (SOP) evolution of the optical propagating beams. As a result, it is rather difficult to evaluate accurately 722, especially in telecommunication fibers where the polarization is not preserved.

In this contribution two simple and effective XPM based techniques that are intrinsically

insensitive to the relationship between the probe and pump SOP's are presented. The proposed techniques allow to evaluate the Kerr coefficient, and hence to recover 722, in common

fiber for telecommunications with an high insensitivity to polarization perturbations. The

first one is based on the use of an Ortho Conjugated Mirror [5], is very simple, accurate and

is useful for low frequencies measurements. The second one uses two orthogonally polarized

uncoherent pumps and can investigate the nonlinear frequency behaviour of the fiber. Both

techniques, moreover, allow to determine the impact of the Kerr nonlinearity in a general

section of a digital transmission optical line.

67
The insensibility to polarization

In a XPM based technique, an intense pump optical field induce, propagating in the optical

fiber, a non linear phase shift on the CW probe. By measuring the peak value A$ of the

induced phase shift, n2 is recovered by the relation:

A$ = 2^P,Le,,6 (1)

where Pp is the pump peak power, Lgj/ the effective length of the fiber, Aej/ the effective core area, A the probe wavelength and the probe SPM contribution has been neglected, b is the polarization parameter and assume an unpredictable value between one third and one,

depending on the relative SOP between probe and pump.

Let's now suppose to have two equal intensity square modulated incorrelated pumps.

Pi and P2, propagating in the Fiber Under Test (FUT) and keeping orthogonal SOP's in

each fiber section. The proof that follows apply also in the case of two orthogonal probes

and a single pump [6]. The two pumps induce a phase shift A#i and A$2 on the probe,

with polarization parameters 61 and 62- The values of 61 and 62 are unknown, as well as the evolution of the polarization states along the fiber, but their sum is constant and equal

to 4/3 whatever it is the input pump and probe SOP and the birefringence of the fibers.

The detected phase shift, A$i + A$25 is therefore polarization independent as well as the

recovered n2.

This can be shown by considering the Poincare sphere shown in Fig. 1, where fields

SOP's in a generic section of the fiber are shown, bi is the weighted sum of the projection of the pump SOP |Pumpi) over the probe SOP |Probe) and its orthogonal SOP

|OppositeProbe). The same apply to the second pump. 61 and 62 are given by

61-1 cos2(|) + isin2(|) and 62 = 1 cos^(^) +

^

sin^(^), (2)

and because a-|-/? = 7r, 61 + 62 = 4/3.

In the following section two techniques that realize this condition are presented, together

with some experimental results.

Experimental setups and results

The first scheme that permits a polarization independent measurement of n2 is shown in

Fig. 2. The scheme is based on the Ortho Conjugated Mirror (OCM), a device that time

reverses the propagation from the point of view of the polarization state [5]. The FUT is inserted in one arm of the Michelson interferometer, fed by the probe source. The square

wave intensity modulated pump beam, powered by an optical amplifier, propagate only in

the FUT. The Kerr action occurs therefore twice: first in the counterpropagating modality

and then, after the reflection on the OCM, also in the copropagating modality. The OCM,

force the SOP of the copropagating probe to be orthogonal to the conterpropagating one in

each section of the fiber and hence the phase shift peak value is polarization independent

[7, 8]. The proof is equivalent to that developed in the previous section [6]. The upper

frequency limit of the pump signal modulation is v/4L^ v being the group velocity and L

68
the fiber length, in order to have also a couterpropagating contribution in the form given

by eq. (1).

If a normal mirror is used instead of an OCM, |Probe) is not anywhere orthogonal to

|OppositeProbe) and the term (61 + 62) remains undetermined. Fig. 3 reports values of

722 /Ae// recovered by measuring the induced phase shift (1) for 10 random input pump

SOP of a DS fiber at lOOKHz. The achieved accuracy is better than 1%. As a comparison,

results obtained with a normal mirror are also reported, showing a dispersion of about 10%.

The second technique proposed can be used to investigate also the nonlinear frequency

response of a fiber. As in the previous technique, the phase shift induced by the pump on

the probe optical beam is measured by an interferometric technique. The insensitiveness to

the polarization is achieved by using as a pump two incoherent and orthogonally polarized

optical beams Pi and P2. The optical circuit is shown in Fig. 4. The pump signal is splitted by a coupler and then recombined by a polarizing beam combiner (PBC). One of

the two beams is delayed by some coherence time. The delay line is then adjusted to obtain a square wave modulated optical pump. The two polarization controllers (PCi,2) permit

to match the right polarization of the PBC and maximize the pump power. The pump

signal is therefore composed by two incoherent optical beam with the same intensity and

orthogonally polarized.

Both pumps copropagate with the probe in the FUT costituting one arm of a MachZehnder interferometer, inducing a non linear phase shift that modify the probe coherent

function [9]. Hence, the intensity (proportional to the phase shift) detected by a balanced

interferometer is the sum of only the two contributions given by the two orthogonal pumps

while the cross contribute is negligible. The detected intensity is polarization independent

because the two pumps keep orthogonal SOP in any section of the fiber. The contribution

of the polarization mode dispersion is negligible because of the small length of the FUT.

At low frequencies (up to lOOKHz) both techniques give the same value of n2, within

the measurements errors. Measurements show a polarization dependence lower than 1%

also for the second technique. Several sections of SMR and DS fibers have been measured

with the pump powered by an ITALTEL AF18-D optical amplifier. FUT length between

200 and 600 meters have been considered. At 1550 nm the low frequency recovered value of

nsMe// is 2.8-10-^°W-^ for SMR fibers and between 4.8 and 6.8-10-^°W-i for DS fibers,

depending on fibers.

References

[1] K. S. Kim, R. H. Stolen, W. A. Red and K. W. Quoi, Optics Letter 19, Feb. 1994

[2] A. Wada, T. O. Tsun and R. Yamauchi, Proc. ECOC'93, paper MoBl.2, Berlin 1993

[3] M. Monerie and Y. Durteste, Electronic Lett. 23, p. 961-963 (1987)

[4] M. Artiglia, B. Ciaramella, B. Sordo, Electronic Lett. 31, p. 1012-1013 (1995).

[5] M. Martinelli, Journal of Modern Optics, 1992, vol. 39, p. 451-455

[6] L. Amato, A. Fellegara, P. Sacchetto, P. Bofii, A. Melloni and M. Martinelli, Optics Letter

21, April 1996

[7] P. Boffi, A. Fellegara and M. Martinelli, Optics Communications, 1996

[8] A. Fellegara, L. Amato, P. Sacchetto, P. Boffi, A. Melloni e M. Martinelli, Proc. OFC '96,

paper ThS4, 288 (1996).

[9] J. T. Manassah, Optics Letter 16, November 1991

69

Interferometric Determination of the Nonlinear Refractive Index n2

ofOptical Fibers

F.Wittl*, J.Vobian**, G.Herchenroder** and W. Dultz**

* Techn. Univ. Darmstadt, Institut fiir Angewandte Physik

** Deutsche Telekom, Technologiezentrum Darmstadt

Abstract

The nonlinear refraction indices of several optical fibers were measured with a Sagnac

interferometer at a wavelength of 1053 nm (NdYLF-laser). The n2-values of the different

fibers were determined to be between 2.0 and 3.3 times 10'^^ m^AV respectively.

Introduction

For high pulse power applications of optical fibers (e.g. soliton transmission) the knowledge

of the nonlinear refractive index is of essential importance. Therefore, the participants of a European COST-Action (COST 241, Characterization of Advanced Fibers for the New

Photonic Network) decided to compare several measurement techniques in order to get the

most reliable n2-values. What we present here, is our contribution to that COST-Action.

For a pulse with the power P(tJ propagating in a medium, the refractive index (n) with

nonlinear correction is commonly described by [e.g. Sto72, Agr89]

« = «o + «2--H- (1)

Agj^is the effective core area of the fundamental fiber mode with the field distribution E(r, (p) and is determined by [e.g. Agr89]

The nonlinear prolongation of the optical path length for a pulse with a central wavelength /I and a peak pulse power P can be represented by an induced phase shift [e.g. Mei95a]

Itt 1 A A,ff

Here, instead of the fiber length L an effective length L^ff has to be used, due to absorption

losses with the absorption coefficients . This effective length is defined as:

l-exp(-«.£)

71
The factor m in (3) is a polarisation parameter. For linearly polarised light w = 1 by

definition of the nonlinear refractive index. For other kinds of polarisation states there are still some discussions about the value of m. If sufficiently long fibers are used (some hundred

meters), polarisation can be regarded as completely "scrambled" and therefore a value of m =

8/9 is now commonly accepted [Eva92],[Cher96].

Measurement technique and experimental setup

Oszilloskop

o

t:

PD

n

J

PI, P2: pulse 1, pulse 2 Gl, G2: glass slabs PMl, PM2: photomultipliers

LI, L2: coupling optics A: vanable attenuator

Pol: polariser DF: detection fiber 01: optical isolator PD: photodiode

PS: polansation controller

Fig.l) Experimental setup

Our measurement technique uses the principle of the Sagnac interferometer (Fig. 1). A similar

setup used by Artiglia et al. and which was also part of the COST 241 project has been

recently published [Art96]. A laser pulse is split into two equal parts by a 50/50-beam splitter.

The two pulses created in this way are counterpropagating through the fiber to be measured.

One of the pulses can be attenuated variably before entering the fiber and therefore induces a smaller nonlinear refractive index than the non diminished one, which passes also the

attenuator, but after being transmitted through the fiber. Due to (3) the induced phase

difference of the two pulses with peak pulse powers P; and is determined by

1 A

m-P,-P, (5)

72
The parts of the pulses passing the beam splitter in the right direction afterwards, are coupled

into a "detection" fiber and due to (5) result in an interference signal detected with a photodiode and an oscilloscope at the end of this fiber. From the measurement of the power

differences - Pjj, necessary to obtain suceeding maxima and minima of the signal, which

means phase differences of AO^ = l-7t{l= 1,2,3...), the effective nonlinearity n2/A^can be

calculated. After measuring the effective core area Ag/f with both refracted near field

technique and transversal offset method [Vob89], the nonlinear refractive index «2 can be

determined.

An Nd:YLF laser was used as light source. We had average powers of 5W-18W, a repetition

rate of 76 Mhz at a wavelength of 1053 nm. The FWHM pulse width of this laser could be

reduced down to 30 ps, but for the sake of better visibility of the interference patterns we only

used FWHM widths of 50 ps - 150 ps. To avoid pulse disturbing feedback effects from the

interferometer to the laser, an optical isolator has to be used.

Results:

Fig. 2 shows some typical measured interference patterns and in Table 1 the determined

values of nonlinear refractive index as well as the effective core areas of some measured

fibers are listed. Table 2 shows n2-values of three of our fibers measured by other

participants of COST-Action at nearly the same wavelength.

a

o

q.

o

A1 AO = 2t

'I \ V

AO = 3;r

J

100 :oo time/ps

fiber type

-20 m_

W

2.1 19.5 Siemens Raman

Fiber

2.5 62.3 Siemens Standard

SMF

3.1 21.1 DC-fiber A, Jydsk

Telekom

3.3 18.5 DC-fiber B, Jydsk

Telekom

3.0 72.4 dep. cladding.

Philips

Table 1 ; Experimental resuhs obtained

for n2 values of some fibres

Fig.2) Maxima and minima of interference of a standard

single mode fiber due to nonlinear phase shift

73
3.0-3.1

@XI nm Fiber type, references

3.0

2.4

1064

1064

1064

DC-fiber A, Jydsk Telekom

[C096], [Mei95]

DC-fiber B, Jydsk Telekom, [C096]

Siemens Standard SMF, [Mei95]

Table 2: Experimental Uj-ydXviQS obtained for the same fibers by other participants of

COST-241 at nearly the same wavelength as our measurements

The experimental results are in good agreement with each other. Also the most drastic

deviation (DC-fiber B, Jydsk Telekom) is between our estimated experimentl error of about

10-15%, which is mostly due to the well known difficulties of determining absolute values of

intensities. Another source of error is that we supposed to have completely scrambled

polarisation, but it is not quite clear whether or not in our measurments (fiber lengths L« 200

m) this can be regarded as true (m=8/9 or m=l). Further measurements are planned in order to

clarify this question.

References:

[Agr89]: Govind P. Agraval: Nonlinear Fiber Optics,

Academic Press Inc., San Diego

[Art96]: M. Artiglia, R. Caponi, F. Cistemino, C. Naddeo, D. Roccato

Opt. Fib. Techn. 2, 75-79 (1996)

[Cher96] : S.V. Chemikov, OFC96, Techn. Digest, ThS5, 289

[Eva92]: S.G. Evangelides, L.F. Mollenauer, J.P. Gordon, N.S. Bergano:

J. Lightwave Technol., Vol. 10 (1992), S.28

[Mei95]: J.Meier: Dissertation, in: Fortschr. Ber.

VDI-Reihe 8, Nr. 443, VDI-Verlag, Diisseldorf 1995

[Mei95a]: J.Meier: 3rd Optical Fiber Measurment Conference

(OFMC95) Luttich (Belgien) 1995, Proceedings

[Sto78]: R.H. Stolen, C.Lin: Phys. Rev. A, At. Mol. Opt.

Phys., Vol. 17(1978), S.1448

[C096]: COST-241 report, S. B. Andreasen, Feb.96

[Vob89]: J. Vobian, G. Herchenroder, EFOC89, Techn. Digest

74
Inverse method for systematic investigations of linear and nonlinear

properties of dispersion-shifted and dispersion-compensating fibers

R. Boness (TU Dresden), S. Unger, J. Kirchhof (IPHT Jena), F. Wittl, J. Vobian (FTZ Darmstadt), K.-F. Klein (FH Giessen-Friedberg), Germany

Abstract

The inverse method allows a defined profile to be calculated making use of prescribed dispersion

curves, cutoff wavelength and field radii. For example, we calculated and verified disperion-compensating fibers (DC fibers) with high compensation ratios (CR) up to 30 with the disadvantage of relatively high attenuation values [1]. In this paper we will present DC fibers with medium CR values and low attenuation values. An interesting new field are dispersion shifted-fibers (DS fibers) with high effective area minimizing nonlinear effects due to low

n2 values.

Theoretical considerations

The effective area Aeff is an important parameter in long-haul fiber transmission systems, above all, in the case of DS fibers. It should be large to minimize nonlinear effects. On the other hand, small Aeff

values of the DC fibers may be advantageous for investigations of nonlinear effects in fibers. In any

case, the dopant concentration should be not too high in order to reduce the attenuation of the fibers under test. We will present an iterative inversion method to optimize the refi^active-index profile and radius parameters with respect to prescribed dispersion and the Petemiann-II field radius values in the vicinity of

1550 run. We assume that Aeff =TUwf is a good approximation of the effective area of the fiber [2]. In

principle, we can investigate any reasonable refractive-index profile, of course within the frame work

of technological limits. In order to determine the corresponding parameters of these special fibers, we have to solve a so-called

inverse problem, which can be done for some prescribed dispersion values at different wavelengths

only using multi-dimensional optimization methods [Ij. We have to minimize a generalized standard

deviation given by

F(x)-X^,(Z)(^,x)-A)^+^,,,(4(x)-A,or+^..2K(^,^)->^/o)' (1)

1=1

with respect to the unknown vector x. X := vector of relative refraction indices Aj {\< i < m) and the fibre radius, thus dim(x) = m+1, in general, m <4.

n' — n'

2A, := —— with n^- refractive index of the outer cladding region in general, 1.4573 «:

Dj := prescribed discrete dispersion values at within 1530 to 1600 nm (1 < / < n with > dim(x)).

A^^:= prescribed cutoff of the LPn-mode, w^-g prescribed PETERMANN-II -radius at AQ=1550nm,

bj := some suitable positive weight factors.

In order to minimize the fijnctional Eq.(l), we make use of the Levenberg-Marquardt algorithm [3] (damped Gauss-Newton method). Our programme is based on a very stable and accurate method for dispersion calculations without any numerical differentiation [1] on the basis of the scalar wave

equation.

75
In formula (1) we assume wfO =5.64 ^m (corresponding to Aeff&lOO |im ) and obtain the profile,

dispersion and field-radii curves presented in in Fig. la,b,c. For comparison in standard dispersionshifted fibers wf ss4.3 |im (Aeff < 60 |.im ), for special DS fibers Aeff values up to 90 \im were

reported [4]. Figures la to Ic document that triple clad fibers (TC fibers) are an optimal solution of

the problem, relatively low dopant concentrations are sufficient, the fiber radii are relatively large

which avoids technological problems.

1350 1400 1450 1500

wavelength nm

Fig I .c: Field radii curves for the profiles (1), (2) and (3).

160

The slope of the dispersion curves at 1550 nm

corresponds to that of standard DS fibers. It

is possible to prescribe other slope values at 1550 nm. The slope may also be zero. The

dispersion and Petermann II field radii curves

for the different profiles in the vicinity of

1550 nm do not differ very much. wfQ is approximately 5.6 \xm, which corr^ponds to an effective area of about 100 ^m . Our

second goal was the development of DC

fibres with a medium compensation ratio (CR=10) and low attenuation values.

In this case we only can prescribe the dispersion values in the vicinity of 1550 nm

but not the Petermann II field radii. Our

inverse method presents the profil curves in Fig. 2a.

76
6

.1 I ^ : ^ 1 012345678: 1

radius jjtn Fig 2. a.: Calculated TC profiles 2A(r) with

constant (1), parabolic (2,4) and linear (3) core segments

14

(4)/

^ A'

i //

...li!>L.z

Petetmann l-radii ;j:<'-

Peiermann ll^radii

wavelength nm

Fig 2.c: Field radii curves for the profiles (1) to (4).

0

-250 I ^ : i i— i J 1300 1350 1400 1450 1500 1550 1600 wavelength nm Fig. 2.b: Chromatic dispersion curves for the

different profiles (1) to (4)

All the profiles (1) to (4) yield CR=10, but

only the profiles (1) to (3) result in a secondorder dispersion compensation. Fig. 2a

shows that the stepwise constant profile

results in the lowest 2A value. The parabolic

profile (4) also results in a low dopant

concentration, but we do not obtain a compensation of the slope.

Fig. 2.C shows that the Petermann II field

radii are smaller than those of Fig. I.e. As a consequence, large n2 values have to be taken

into account.

Measuring procedures and practical results

First experiments were undertaken to fabricate and to analyze the features of DC and DS fibers which

were fabricated by MCVD. A general problem of the fiber preparation is the realization of a high refi-active inde.x in the core together with small fiber-core values, especially in the case of DC fibers. The

germanium content of the core is determined rather by diffusion than by the chemistry of incorporation

during the deposition because of the strong increase of the diffusion coefficient of germanium with the concentration [1]. Fig. 3 .a,b show the TC profiles of the DS and DC fibers measured at the preforms together with the

corresponding ESI curves. The mode-field diameter wf and consequently the effective area was determined by making use of the transverse offset-method. From the obtained autocorrelation MFD wf was calculated using the formula of Calzavara, Coppa, Di Vita [5], The chromatic dispersion measurements were performed with the aid of a raman fiber laser with a mode-locked Nd:YLF laser used as a pump source and a Mach-Zehnder interferometer for measurements on short fiber pieces [1]. The measurement of the nonlinear Kerr coefficient n2 was performed making use of the Sagnac inter- ferometer principle [6]. Thereby a pulse of a mode-locked Nd:YLF laser is split into two parts which

are passing through the fiber in opposite directions and have different power values. For detection

the two pulse parts are superimposed again. From the interference pattern the nonlinear phase shift and

consequently the nonlinear refractive index n2 can be calculated.

77
Fig. 3a TC profile of the DS fiber Fig. 3b TC profile of the DC fiber

The measured effective area of the DS fiber yields 74 jim^. Values up to 100 ^m^ can be verified. . The measured effective area of the DC fiber is 23.8 [im^, in good agreement with the theoretical value

(Fig. 2.C). The Kerr coefficient of the DS fiber was determined at « 2.5* 10'^*^ m^W"^. of the DC fiber was not measurable due to the stimulated Raman scattering effect.In any case, the wavelength was 1550 nm.

Figs. 4 and 5 present the measured dispersion curves of the DS fiber and the DC fiber, respectively.

29 - ps/iifn*kn»

10 h

ISBO I see

wavcleitgtU i

wavclcii^jlh m ntn -29

Fig. 4 Dispersion curve of the DS fiber Fig. 5 Dispersion curve of the DC fiber

References

[1] R. BonelJ, W. Nowak, J. Vobian. S. linger, J. Kirchhof Tailoring of dispersioncompensation fibres with high compensation ratios up to 30. Pixre Appl. Opt. 5 (1995), p. 333

[2] P. Nouchi, P. Sansonetti, J. von Wirth, B. Perrin, Alcatel Alsthom Recherche (Marcoussis,

P.): Large Effective Area Dispersion Shifted Fibre for Long Haul and High Bit-Rate

Transmission. EFOC&N (Brighton), June 1995, pp. 101-103

[3) J. M.Ortega, C.Rheinboldt.: Iterative Solution of nonlinear Equations in several Variables.

New York. Academic Press 1 970

[4] Y. Liu, A. Joe Antos, M. A. Newhouse: Large effective area dispersion-shifted fibers with

dual-ring inde.x profiles. OFC '96 Technical Digest. Section WK 15, pp. 165-166

[5] J. Vobian. G. Herchenroder, E. Unterseher: Comparison of mode-field diameters of normal and advanced

SM- fibers measured with the transverse offset-method with calculations based on refractive index measurements. EFOC/LAN 89 (Amsterdam). June 1989. pp. 232

[6] F. Wittl, J. Vobian. G. Herchenroder, W. Dultz, T. Tschudi: Interferometrische Bestimmung des nichtUne- aren Brechungsinde.x n2 optischer Fasern. 97. Jahrestagung der DGaO, Mai 1996. Neuchatel/Schweiz

78
Effective Area Measurement Comparison Between

Direct Far Field Scan and

Variable Aperture Method in the Far Field

Michael J. Hackert and M. Ajaaz Zainul

Coming, Inc. Corning, New York 14831

Summary

A key parameter in determining the impact of non-linearities such as self phase modulation (SPM) and four wave

mixing (FWM) in high speed, optically amplified systems is the effective area (Agff) of the fiber. This paper

describes modeling and experimental results for measuring \{f using direct far field scanning and the variable

aperture method in the far field (VAMFF). The modeling and experimental results confirm that both measurement

methods give equivalent results. However, far field data (by either method) must be collected to at least 0.4 numerical aperture (NA), but preferably 0.5 NA or larger to minimize the error resulting from truncation of the data

collection. The impact of this truncation is evaluated in this work.

Introduction

As the state-of-the-art in fiber optic systems has pushed to faster speeds and longer lengths, the impact of non-linear

effects such as self phase modulation (SPM) and four wave mixing (FWM) has increased in importance.

Fortunately, these effects can be effectively modeled using the non-linear coefficient n2/Aeff (non-linear refractive index / effective area).' Variation in the effect of non-linearities from fiber to fiber is controlled by effective area

(Aeff). Thus, Aeff is a key parameter to be monitored.

Aeff can be measured by two techniques^ - direct measurement of the far field pattern and the variable aperture

method in the far field (VAMFF). These are the same techniques used for measuring mode field diameter (MFD).

In this paper, a comparison is made between the two methods for measuring Aeff. The comparison has two parts.

First, the problem is approached analytically using modeling. The percent disagreement between the two methods is predicted. The second part summarizes the experimental results. Also, the impact of the maximum NA of the VAMFF system is evaluated. Similar comparisons for MFD were also made to confirm the analysis methods and

experimental set ups.

Modeling Results

The simulations started with direct far field scan data. Power as a function of angle was gathered as described in EIA/TIA-455-164A^ . An example of 3 types of fiber profiles were included in the analysis - standard single-mode

fiber (Coming's SMF 28TM), standard dispersion shifted (DS) single-mode fiber (Coming's SMF/DSTM), and large

effective area dispersion shifted fiber. Readings were obtained up to a maximum available numerical aperture (NA)

of 0.3338 for this particular far field scanner.

Aeff and MFD were then found using the direct far field scan formulas (1), (2) and (3) given below:

MFD = 2w 0 =

(1)

79
l2

where

[f E(r)' rdr]

A..= In

^=L (2) j^^E(r)'rdr

E(r) = Bo j^"[P„(e)]'" Jo(rPsine)sin 2ede (3) Pm is the far field power distribution,

9 is the subtended half angle in the far field , r is the radial distance from the fiber's core ,

(3 = InJX,

Jo is the zeroeth order Bessel function of the first kind, and

E (r) is the near field's electric field.

The next step was to select a definite set of values for the apertures on an aperture wheel for a VAMFF system. The

corresponding powers for the variable aperture technique were then simulated using the same power versus angle

data. MFD was then determined from the simulated VAMFF data using formulas (5) and (6) utilizing a trapezoidal

fit. MFD for the VAMFF is given by

V2X

where

MFD = 2Wo =^^|^jja(e)sin2ede] (5)

a(e)= 1- PQ)

V P max

(6)

P (9) is the power transmitted through each aperture, Pmax is the power transmitted through the largest aperture, and

9 is the subtended half angle of each aperture from the fiber end.

Use of formulas (7) and (8) below together with formula (2) above allowed Agff to be determined from the modeled

VAMFF data.

P dp

where

E(r)= EoJJ°[Pppf'Jo(rp)dp

PvAMFF is the power through each aperture on the VAMFF,

p = (27i/>.) sin (9), PpF is the equivalent average direct far field power,

J,) is the zeroeth order Bessel function of the first kind, and

r is the radial distance from the fiber's core.

(8)

The aperture wheel in Table 1 with a maximum numerical aperture of 0.3329 was selected for the VAMFF system.

Table 2 summarizes the subsequent results of the modeling. The results indicate there is good correlation between

the two systems for measuring Agff and MFD. The largest disagreement was on DS and Large Aeff DS fibers. This is likely attributable to the limitation set by using a maximum numerical aperture of 0.3329. If the maximum numerical

aperture is increased, the predicted differences between the direct far field measurement and VAMFF on DS fibers should reduce quite markedly.

80
Test Results

In order to empirically confirm the conclusions from the modeling work, a direct far field scanner was set up

according to FOTP 164^. It was designed to have a maximum scan of greater than 0.6 NA and a dynamic range of

greater than 55 dB. Tests confirmed that its angular accuracy was better than ±0.0005 radian over its full scan angle

and that its detector was linear to approximately 5% over its dynamic range.

The experiment consisted of measuring a sample of 9 fibers including standard single-mode fiber, standard

dispersion shifted (DS) single-mode fiber, and large effective area dispersion shifted fiber. The fibers were chosen

to span a wide range of \{f values. The samples were measured on both the direct far field scanner and the VAMIT

measurement system at both 1310 and 1550 nm. The maximum angle for the far field scans was 0.4 NA for the standard single-mode fiber and 0.6 for the other fibers. The fibers were measured 3 times each and the data used to calculate both MFD and Aeff. The MFD values agreed within approximately 0.05 |im (< 1%) confirming the accuracy of the two measurement systems. The average Aeff values are given in Table 3. The average bias for the 9

samples is also given (-0.28 and -1.81 |im^ at 1310 and 1550 nm, respectively). These results are consistent with the

results from the modeling and show that both measurement methods have equivalent ability to measure Aeff.

The last evaluation performed was to determine the impact of restricting the NA used in the calculation. The impact

is shown in Figures 1 and 2 for Aeff and MFD, respectively. The reladve difference in each parameter compared to the value at 0.6 NA is plotted versus the maximum NA of the calculation. It is evident from the analysis that the impact is relatively negligible as long as the NA is greater than 0.4. Of course, an NA of 0.5 or 0.6 is more

desirable.

Conclusions

Aeff is a critical parameter in determining the non-linear performance of high speed, amplified fiber optic systems.

Both mathematical modeling and empirical test data have shown that both the direct far field scan data and the variable aperture method in the far field (VAMFF) methods of measuring Aeff give equivalent results as long as the data included in the analysis includes sufficient angular extent. A minimum of 0.4 NA is required, but 0.5 NA or

greater is desirable.

References

1. "Comparison of Non-linear Coefficient of Optical Fibers at 1550 nm," Valeria L. da Silva, et. al., ThA4, OFC

1996.

2. EIA/nA-455-164A "Single-mode Fiber, Measurement of Mode Field Diameter by Far Field Scanning,"

EIA/TIA-455-167A Mode Field Diameter Method in the Far Field," draft TIA/EIA-455-132 "Measurement of

the Effective Area of Single-mode Fiber."

0.0139 0.0218 0.0264 0.0350 0.0433

0.0494 0.0641 0.0785 0.0930 0.1069

0.1211 0.1364 0.1521 0.1619 0.1770

0.1921 0.2266 0.2552 0.2928 0.3329

Table 1 - The values of the numerical apertures used on the aperture wheel on the VAMFF system that produced the modeling results in Table 2.

81
Aeff MFD

(Um)

Direct Far

Field

VAMFF Percent

Difference

Direct Far

Field

VAMFF Percent

Difference

Dispersion

Shifted Fiber 44.00 45.02 2.32 7.693 7.859 2.17 Large A^ff DS

82.64 84.29 2 10.07 10.25 1.8 Standard SMF

74.45 74.65 0.27 10.064 10.229 1.6

Table 2 - 1550 nm modeling results using the VAMFF system aperture wheel in Table 1.

Aeir

FIBER ID

AVERAGE

VAMFF

1310 nm 1550 nm

AVERAGE

Direct FF

1310 nm 1550 nm

Delta

(lim')

1310 nm 1550 nm

Delta

(Percent)

1310 nm 1550 nm

DS#1 28.83 43.12 28.82 43.53 0.01 -0.40 0.0% -0.9%

DS #2 32.27 51.16 32.29 51.31 -0.02 -0.16 -0.1% -0.3%

Large Agff DS 1 47.30 84.14 47.27 85.37 0.04 -1.24 0.1% -1.4%

Large Agff DS 2 34.39 59.46 34.56 61.68 -0.17 -2.22 -0.5% -3.6%

Large Agff DS 3 44.70 68.91 37.97 72.24 6.72 -3.33 17.7% -4.6%

Std SMF #1 63.52 78.87 65.55 81.54 -2.03 -2.67 -3.1% -3.3%

Std SMF #2 64.65 81.11 66.96 83.25 -2.31 -2.14 -3.4% -2.6%

Std SMF #3 66.34 83.26 68.93 85.22 -2.59 -1.97 -3.8% -2.3%

Std SMF #4 66.52 83.59 68.68 85.76 -2.16 -2.17 -3.1% -2.5%

Average

Delta

-0.28 -1.81 0.4% -2.4%

Table 3 - Agff comparison between Direct Far Field (FF) and VAMFF.

Fig. 1 Maximum NA's impact on Aeff bias Fig. 2 Maximum NA's impact on MFD bias between direct far field and VAMFF between direct far field and VAMFF

82
Characterizing Components for High Speed Data Interconnects

David Smith

Honeywell, Micro Switch

Richardson, Texas , USA

dsmith@micro.honeywell.com

Introduction

The rapidly growing use of optical fiber in high speed LANs and data interconnects has placed new demands on

technologies required to support them. Although fiber was employed in 10 Mb/s Ethernet links, it was not used near

the limits of its performance capabilities. Even in lOOMB/s FDD!, where the sources were typically 1300nm LEDs,

the fiber was not really the component that limited link performance.

At data rates around 0.5 GB/s, 1300 nm LEDs still provide acceptable performance, although the combination of slow edge speeds, low coupled power and large spectral width placed these components close to the limits of their performance. At data rates beyond ATM 622 MB/s ,there is no feasible alternative to using a semiconductor laser and in fact, many of the standards permit lower cost 850 nm lasers and LEDs to be used at data rates lower than 1 GB/s ( e.g. Fiber Channel 531 MB/s, 265 MB/s and ATM 155 MB/s). However, even

lasers do not solve all problems, and at rates above 1 GB/s, the modal bandwidth of multimode fiber begins to limit the transmission distances even at 1300nm.

For cost reasons, AlGaAs CD lasers have generally been employed at speeds up to 1 GB/s, but a new type

of laser, the Vertical Cavity Surface Emittmg Laser (VCSEL) is showing great promise as the source of the future because it has the potential of meeting all the system requirements and simultaneously being low cost, high speed,

low coherence and can be manufactured as an array for parallel optical buses. It is anticipated that VCSELs may

eventually replace other types of sources. Although there is much published in the literature concerning VCSELs,

little has been published about the characteristics and performance of these devices in typical data communication

links, where their first use as commercial sources is already emerging. This paper will address some of the issues surrounding the characterization of optical sources, in particular VCSELs, for high speed data communications.

Characterization of VCSEL sources. VCSEL Structure The VCSEL structure described in this paper is a high reliability, multimode, gain guided device with

proton bombarded carrier confinement implants as described in [1,2]. The device, intended for commercial LAN

use, has an implant inner diameter of 20 microns and an aperture of 15 microns. This provides the best balance

between modulation rate, threshold current , series resistance and spectral width. However, apertures up to 30

microns are also employed. All these devices are designed to be single longitudinal mode but multi-transverse mode. The simultaneous lasing in multiple transverse modes ensures low source coherence, greatly minimizing

interference effects like modal noise compared with CD lasers. [3]

Modal Noise.

Modal noise [4] occurs when a coherent source generates time-varying, modal interference patterns

(speckle) in a multimode fiber and these speckles encounter a mode selective loss such as an offset lossy connector

joint. The ensuing degradation of the optical signal-to-noise ratio can generate error floors if the source has adequate coherence, and the loss is large enough. Since these amplitude fluctuations can have a large spectral

component at very low frequencies (drift) in a real data link, the statistics are effectively non-stationary, making

their effect difficult to predict. Consequently a test methodology, based on [5] is under development in TIA/EIA

FO-6.5 [6] to enable the accurate measurement of the modal noise penalty created by mode selective loss. The

suitability of the source is assessed by measuring its BER performance in a test path containing 3 separate 1 dB

mode selective loss (MSL) points placed close to the laser and followed by 200m of multimode fiber. This BER is compared to the BER in a reference path with no MSL, but otherwise identical. The statistics are made stationary over the duration of the measurement by employing a "fiber shaker" which agitates the fiber and thoroughly

83
explores the speckle states. The amount of sensitivity penalty

tolerated is specified by the data link standard, but typically 0.5 dB is used for 62.5 /125 micron fiber. Due to their low coherence,

VCSELs show significantly lower penalties than CD lasers for an

equivalent amount of MSL (Fig. 1). This raises the possibility that unlike CD lasers, VCSELs will not require any testing for modal

noise since they are have adequately low coherence by design and

the link specifications limit total connector loss to 3 dB.

Although, to the author's knowledge, there is not a single instance of a short wavelength link employing CD lasers ever having

failed due to modal noise in the field, this will help alleviate the impression that modal noise is a problem for laser links.

BER vs Average Power

•23 -22 -21 -20 -19 -18 -17

Bit error

rate

Fig 1 . Modal Noise Comparison of 850 nm VCSEL

and 820 nm CD laser with 6dB MSL.

Modulation Characteristics

The modulation characteristics of lasers are critical for

their use in data communication links. These characteristics include such parameters as rise time, fall time, jitter, small signal

bandwidth, duty cycle distortion and turn-on delay. The exact values of these parameters are precisely defined in the appropriate standards for any given protocol but in practice the suitability of the laser is determined by

superimposing its 'eye' pattern on a mask defined in [7]. If VCSELs are required to support data rates beyond 1 GB/s it is essential that their modulation properties are well understood, and characterized. Fig 2 shows the schematic of a 'typical' optical

pulse from a VCSEL. Most VCSELs have a 0-80% rise time and 100%-20% fall time that is extremely fast. Properly designed gain guided , multimode VCSELs,

routinely have rise times <60 psec and fall times < 100

psec when measured from the 20 - 80% points ( Fig 3). These fast edge speeds are consistent with the Fourier

Transform of the small signal modulation bandwidths

observed (Fig 4) when the VCSEL is biased above

threshold and close to its peak operating current (5-10

mA). Fig 5 shows that these VCSELs can support data

rates at least up to 3 GB/s sec which should be adequate

for data link communications for many years in the future.

Fig 2. Schematic of 1 nsec VCSEL pulse showing

characteristic features.

As can be seen in fig 2, however , the final 20% of the pulse amplitude can exhibit anomalous tailing effects . The duration of these effects is usually independent of the fast edge speed but may extend to 1.5 nsec. Different devices adjacent on the wafer can show different behavior. At

one extreme some devices exhibit an overshoot followed by a slow

decay, at the other extreme, a slow rise. In general the behavior can be

anywhere between these extremes and can vary with bias conditions on

any one laser. A. Valle et al. [8] have shown that when a lateral diffusion term

is added to the rate equations, spatial hole burning and carrier diffusion

create a mechanism for strong competition between modes in particular

if the modes are strongly overlapped, often leading to dominance of a

single mode. They also point out that if the modes are sufficiently

spatially independent, the may lase independently although they are competing for the same total gain. On turn-off, spatial hole burning can also induce a carrier gradient which causes carriers to fall into the region which

VCSEL rising and faltins edges

Fig 3. VCSEL rise/fall times

84
Freq Response 473(1906)20/15

-15

-18

-21

-24

-27

-30

3 \.;

ma

i

V L 12 na

Smau ^26mi 1

aooo 10000 12000 freq (MHz)14000 16000 1(000 20000

Fig 4. Small Signal Frequency Response

has already been turned off, and cause it to

rise briefly above threshold again. This

would explain the off-state bounce that we

observe. This theory does not explain all the effects on the rising edge however.

Rising edge speed is affected by

differential turn on delay of different modes, and diffusion effects caused by

lateral inhomogeneities as the carrier density reaches equilibrium. A feature unique to multi-transverse mode lasers where there is some spatial independence

between modes, is that there appear to be

regions of the laser which may remain

3 GB/s VCSEL eye

50psec/div

Fig 5. 3 GB/s "Eye"

sub-threshold when other regions are already lasing, even though the entire region is being pumped. This means that the non-lasing regions will have

long carrier recombination lifetimes since they are dominated by spontaneous

recombination.

Evidence for this has been observed by examining the emissions

from VCSELs in the plane of the junction. Test structures were made by

surrounding the VCSEL with a secondary VCSEL structure and using it as a photodetector. Fig 6 is a plot of the detected photocurrent vs. laser bias current from this structure. This shows that the spontaneous emission does

not entirely clamp at threshold as expected from simple laser theory. In fact it continues to rise, although there is a clear discontinuity in the curve as each

mode reaches its own unique threshold, suggesting that at least locally, the spontaneous emission is pinned as expected from conventional laser theory.

However, even at 10 ma the spontaneous emission has not fiilly saturated. These sub-threshold regions act like parasitic capacitance hanging on the

lasing regions. Since this effect implies that the quasi-Fermi levels are not uniformly clamped across the active region, this can only be true if there is enough lateral resistance to permit a difference in potential difference across

the device. The combination of the capacitance and resistance may well resuh in a slow turn-on due to the RC time

constants alone, but also the carrier inhomogeneity induced by this mechanism will cause carriers to diffuse from the non-lasing

regions into the lasing regions presumably

limited by the carrier lifetime which is of the order of the 1.5 nsec tail that we observe. Given

the complexity of the VCSEL mode pattern, this

is an extremely difficult effect to model

precisely.

None of these mechanisms appear to explain why some lasers show a significant

overshoot on the rising edge even after the

relaxation resonance has been filtered. However,

the inclusion of modal "shadowing" ( i.e. light not being emitted due to lasing under the metal

contact) in the rate equations shows that such an

Fig 6. Spontaneous Emission effect is possible. Lam [9] has shown in a study

Vcsel 603 Integrated Monitor @ 20 Deg. C

3000

2000

1500

500

Spontanvout

Emiftton ^ .^_-<-^^Output

Monttor jr^ Powtr

Current yz5->- ^Stimulated Emitslon

125.0

100.0 t

75.0 Z

2 3 4 5 10 11 12 13 14 15

85
Simulated pulse responses with

different modal combinations

Overshoot . Slow modes removed from total

Flat Response. Total Power

Slow Rise. Fast mode removed from total

Filter Bandwidth 1.5 GHz

0.0 0,5 1.0 Time (ns)1.5

of pulse characteristics in quantum-well edge

emitting lasers that even if the total photon output

reaches equilibrium rapidly, the individual modes of

a multimode laser may take many times longer to reach equilibrium. Some modes initially overshoot

beyond their equilibrium, others undershoot and

climb to equilibrium slowly. Clearly, these two

behaviors must compensate each other when the modes are added to give the total power. However,

if some of the modal power is occluded, then the compensation will not be perfect and the result may

be a photon output which either overshoots or

exhibits a slow rise depending on which modes are blocked. Fig 7 shows the result of a simplified rate equation simulation which limits the number of

lasing modes to three. The three curve families show

the total power, an instance where the slower modes

Fig 7. Effect of selective modal transmission

were shadowed producing overshoot, and an example where the

faster mode was shadowed, producing a slow rise. In each case the

plot contains both the unfiltered waveform and a version which

has been filtered with typical system filtering that one would

expect in a gigabit system application. This example is intended to be illustrative only and exaggerates the effects. Fig 8 shows some

actual results with a VCSEL which showed these effects clearly

and which started lasing in a ring near the contact pad where

shadowing is likely to occur. The different traces represent

different bias currents and the traces are scaled and overlaid to ease comparison of the effects. In this case the fastest pulse

corresponded to the highest drive current. The impedance of the VCSEL was 20 Ohms and an external series resistor of 30 Ohms

was added to obtain impedance matching. The VCSEL impedance

only varied by 5% throughout the measurement. A more detailed analysis of this mechanism is to be presented in a fiiture paper.

Comparison of tait:ing (effects on risingiedge of ycsEL Pulses 1

\\ \ lth+5ma

\ \ : ' ith+3ma ;

Ith+1ma

2 nsec/div Ipk-pk = 5ina

RXfaiandwidthSOO MHz lth=>3ma

- . • : • -

Fig 8. Rising edge distortion effects

Subthreshold

bias level

» '-1 .9mA -—-1.4mA

0.0OE*00 S.OOE+OS 1.00E+09 1.50E+09

Bit Rate Fig. 9 Turn-on delay vs bit rate for various

bias levels

The importance of understanding this mechanism

lies in the fact that although they are fast, multimode

VCSELs should exhibit some slow tailing due to their transverse structure. It is important that characterization of

rise time is not fooled by these shadowing effects and the

laser is designed to eliminate this effect.

Systematic jitter.

The main source ofjitter in any laser, other than

parasitics, is turn on delay. Fig 9 shows the measured rising

edge turn-on delay vs. bit rate for a VCSEL using a 5 mA

pk-pk modulation current pulse under different bias

conditions. The definition of turn-on delay in this case was

the delay of the rising edge measured to the pulse midpoint

(where zero delay was defined as the delay observed for a

86
Fig 10. Normalized Tlireshold Current vs Temperature

pulse biased far above threshold) for a

101010... data stream at the system

clock rate. As expected, turn on delay

is reduced as one increases the bias

level closer to threshold. [10] The

turn-on delay reduces as bit rate increases for constant bias conditions. However, this result is an artifact of the measurement technique where the data

is an NRZ 101010... stream. The pulse

must be adequately long to allow the

carrier density to reach its equilibrium

"off' state, when the laser is turned off For short intervals between pulses the

carrier density may not reach

equilibrium and pattern dependent

effects may be present. When a laser is being characterized for turn on delay

jitter, it is essential that either pseudo

random data is used or pulses of

adequate length are used if a PRBS is not available to avoid this effect. This is particularly true if the VCSEL is being modulated from zero-bias [1 1], or when trying to assess the effects ofjitter on the BER when running the VCSEL with no feedback control [12]

40 50 60 Temperature °C

DC_591A.s

7(19/95

477.5tiV

Threshold temperature sensitivity and open loop bias control. Like all lasers, VCSELs have threshold currents which vary with temperature. However, because the VCSEL is a single longitudinal mode laser whose lasing wavelength can be placed anywhere on the gain peak, the reduction in gain with temperature can be offset by the change in the peak wavelength of the gain peak with

temperature. This allows the designer to select a wavelength at which the threshold shows very little change with

temperature. Fig 10 shows the normalized change in threshold with temperature for many devices scattered over a

wafer. It can be seen that a significant range of devices exhibit a very low sensitivity of threshold to temperature over what would be described as a typical operating temperature range in a LAN application. This allows selected VCSELs to be operated in an open loop configuration with no feedback or monitoring circuitry. The device can be

biased at the lowest threshold in the desired temperature

range, and the amount of drift that can be tolerated would

depend on what level of turn-on delay can be tolerated by

the system. [12] Fig 9 shows that for systems around 1 Gb/s , where the peak to peak systematic jitter limit is around 180 psec, this approach is feasible if the VCSEL

bias is kept within 0.9 mA below threshold. For lower bit rates this range extends further. This method could be

particularly useful where a very simple architecture is required, for example driving a VCSEL directly from a

logic gate, where eye safety levels are of minor

importance, and where there is no space to include

monitor detectors to control the bias point for example in high density, low cost back plane applications. Devices

which do not meet these requirements, particularly at high

data rates can be used with a monitor diode to establish a control loop.

Fig 1 1 . Effects of Incomplete fiber coupling on VCSEL modulation, coupling Loss

Into MM fiber

, ^ 0.5 dB

2.4 dB

161.

Coupling the source to a fiber. An equally important part of using a laser source is the ability to couple it to an optical fiber. VCSELs

87
are useful devices in this respect. In fact multimode VCSELs, which commonly range from 15 micron diameter to 25 micron diameter active apertures are particularly easy to couple to optical fibers. They do not require high

magnification lenses , they have circular beam geometry, and have relatively low NA around 0.1 - 0.15. This would

allow a fiber to be coupled directly to a VCSEL, but since all commercial transceivers are connectorized , a lens

will generally be employed for reliability purposes to translate the effective source point to a place where there is no

risk of damaging the VCSEL with the fiber.

The optimal coupling configuration of a VCSEL or any laser to a multimode fiber is to launch a spot centered on the fiber core, with a diameter less than half the fiber core diameter. This will achieve a number of

goals. It will eliminate modal noise resulting from connector loss .[13] It will eliminate most connector loss in a

link. [13] It will enhance the bandwidth of multimode fiber by reducing differential mode delay with most FDD!

compliant fiber. It will permit use of either 62.5/125 micron and 50/125 micron graded index MM fiber with no

coupling loss. It will prevent variations in the VCSEL numerical aperture from affecting coupled power. Fig 1

shows what happens to the pulse when a VCSEL overfills a fiber at launch (up to 6 dB coupling loss). If the fiber is overfilled, the coupled power varies with time due to an increase in the NA of the VCSEL beam as a result of

thermal lensing and mode focusing. [14,15] The effect occurs primarily at the thermal time constant of the chip, and

has a much smaller effect for more reasonable coupling losses. Fig 1 1 shows that it is negligible at 1 dB coupling

loss. This type of distortion would cause problems in high speed systems only if one attempted to set up the AC

power levels by measuring the DC L-I characteristics of a device. This is not a trivial issue however, since for manufacturing purposes, it is desirable to avoid AC testing. Clearly it is best to avoid this problem at source and

ensure that there is no mode selective coupling loss at the connector.

Conclusion.

VCSELs are an ideal source for future high speed LAN applications due to their high speed, good

reliability, low coherence, and the ability to use them in array configurations and other high density applications

with no monitor circuit and no feedback control. However, being multi-transverse mode devices , they have

attributes that require careful management and characterization to ensure satisfactory performance. This is a

relatively new, but fast growing technology and one would expect to see these characteristics designed out of the devices in a manner similar the design improvements of edge emitting lasers.

References

1 . R.A.Morgan et al., IEEE Photonics Technology Letters, Vol 7, No 5, May 1995, p441.

2. J.K.Guenter et al. Proceedings of the SPIE, Vol. 2683.

3. Hahn, K.M., et al. Electronic Letters, Vol. 29, No. 16, Aug 5, 1993 pp. 1482-1483.

4. R.E.Epworth, Technical Digest, Fourth European Conference on Optical Communication, Genoa (1978),

pp 492-501.

5. R.J.S Bates et al.. Optical and Quantum Electronics, vol 27, pp. 203-224,1995.

6. Draft Worging Group Document describing the measurement technique is available. Contact the author by

e-mail for a copy.

7. "Optical Interfaces for Equipments and systems relating to the synchronous Digital Hierarchy." ITU-T

Recommendation G.957 1993.

8. A. Valle, J. Sarma, K.A. Shore, IEEE Journal of Quantum Electronics, QE-31, No 8, Aug. 1995, pp

1423-1431.

9. Y.Lam, J.P.Loehr, J.Singh, IEEE Journal of Quantum Electronics, QE-28, No. 5, May 1992. 10. Chinone et al., IEEE J.Quantum Electronics, QE-10, 81 (1974).

11. D.M. Currer, K.Y.Lau, IEEE Photon. Techol. Lett., Vol. 7, pp. 4-6, 1995. 12. K.Obermann et al., IEEE Photon. Techol. Lett., Vol. 8, No.l pp. 31-33, 1995. 13. T.H.Wood, L.A.Ewell, IEEE J. Lightwave Tech. Vol LT-4, No. 4, April 1986. 14. Zhiao,Y., Mclnerny, J.G., Morgan, R.A., Optical Engineering, Vol. 33, No. 12, Dec 1994

15. Zhang, J-P., Petermann, K., lEE Proc Optoelectron., Vol. 142, No.l, Feb 1995

88
Modal Optical Coupling Efficiency Measurements and

Modeling for Fiber Interconnections

J. Sutherland, G. George and J. P. Krusius

Cornell University, School of Electrical Engineering, Ithaca NY 14853

Abstract - Measurements and modeling of modal optical coupling efficiency have been performed on

fiber-to-fiber interconnection links. Modeling results have been validated with measurements of single

mode to multimode fiber coupling for axial separations and lateral misalignment. The model agrees with

power-normalized coupling efficiency measurements under these conditions to within 15% for large axial

separations. The modeling approach is suitable for fiber-to-fiber, laser diode-to-fiber, fiber-to-rectangular

waveguide interconnections.

1. Introduction

Typical optical links employ multiple fiber-to-fiber interconnections. Minimization of coupling losses at each interconnection leads to improved bit error rate (BER) performance and reduced power dissipation at sources and detectors. Multimode fiber-to-fiber coupling has been modelled using ray tracing techniques

[1-3], but these methods are unable to account for variations in launching fiber mode power amplitude

and phase distribution, and they cannot predict the modal energy distribution in the receiving fiber. We

have developed a mode-based optical coupling model which predicts coupling efficiencies for multimode

fiber-to-fiber interconnections with axial separations and lateral misalignments [4,5]. The approach

requires prior knowledge of the modal power distribution in the launching fiber. Since the modal power

distribution is general difficult to measure in multimode fibers, this paper presents coupling efficiency

measurements and modeling for a configuration where the source modal power distribution is known:

single-mode to multimode fiber coupling. The model is also applicable to emerging low-cost laser diodeto-fiber couplers where alignment tolerances are relaxed by coupling into large diameter multimode fibers

[6,7], and fiber-to-rectangular waveguide couplers.

2. Coupling Efficiency Measurements

Hber type core dla.

HP8153A

Lightwave

Multimeter

Single-mode

Multimode

Source fiber: 2230 in length of single- mode fiber on spool

8.3 |Llm 50 |am

cladding dia. 125 mmAn

125 mm

0.36%

3.0%

Coupling efficiency measurements were performed using the experimental setup shown in Fig. 1 . Light was launched into a 2230 m long single-mode optical fiber from an HP1552SM 1322 nm laser source. The single-mode fiber had a nominal core diameter of 8.3 |im and a step-index profile with An = 0.36%.

These fiber parameters ensure

population of only the lowest

order mode, with cladding and

unguided modes stripped due to the fiber's length [8]. Light was

coupled from the single-mode

fiber into a 50 p.m core diameter

multimode fiber with 125 |J.m

cladding diameter and a step-index

profile with An = 3%. The

multimode fiber was 50 m long to help strip out cladding and

unguided modes. The far end of

the multimode fiber was attached

to an HP 81532A optical detector,

which provides power measurements accurate to ±4%. Fig.

fiber coupling

site

Receiving fiber: 50 m

length of multimode

fiber on spool

|1.33 dB

1

HP81S32A

Optical Detector HP81352SM

1322 nm SM

Laser Source

xyz-position

stages, accurate

to 0.5 |lm

1 . Experimental setup used to measure optical coupling

efficiency for 1322 nm light coupled from a single-mode fiber into a multimode fiber.

89
Axial and lateral misalignment of the two fibers was controlled using two Newport 561 XYZmicropositioners. Each micropositioner has been characterized for repeatable positioning accurate to 0.5

|j.m. Fiber end faces were prepared using a Fujikura CT-04 high precision fiber cleaver. No end face

polishing was performed, but all fiber ends were optically inspected for surface defects prior to measurement. Averaged measurement results for axial separations and lateral misalignments are shown

in Figs. 2 and 3, respectively. Both figures illustrate the effects of reflections at short axial

misalignments.

o.2 r I I I I , , I , I I I I , I , , I I 1 0 100 200 300 400 500

Axial separation (|iin)

Fig. 2. Measured single-mode to multimode fiber coupling efficiency for axial separations.

0.8 ^ ' ' ' '

I

' M ' I

Lateral misalignment (|im)

Fig. 3. Measured single-mode to multimode fiber coupling efficiency for

lateral misalignment at various axial separations.

90
2. Optical Coupling Modeling

The single-mode to multimode fiber coupling configuration shown in Fig. 1 was simulated using our modal coupling model. The simulation program computed source fiber mode, diffracted the mode onto

the receiving fiber face, and computed the numerical inner product of source modes onto the receiving

fiber modes. Derivations related to source and receiving fiber mode structures, efficient diffraction of source modes, and overlap integral computation have been presented previously [4,5]. The coupling

model includes consideration of reflections at the source and receiving fiber faces. Computation of

coupling efficiency for each axial separation and lateral misalignment case required approximately two

minutes of CPU time on an HP-700 series workstation.

Axial separation (^m)

Fig. 4. Simulated single-mode to multimode fiber coupling efficiency for axial separations.

20 30 40

Lateral misalignment (|im)

Fig. 5. Simulated single-mode to multimode fiber coupling efficiency for

lateral misalignment at various axial separations.

91
Figs 4 and 5 show comparisons of measurement and simulation for axial separations and lateral misalignments, respectively. Reflection effects are evident in Fig. 4 for small axial separations. The

limited number of data points leads to an aliasing effect that can be corrected with additional data points.

Comparison of measurement and simulated coupling efficiency plots for axial separations and lateral misalignment shows agreement in general shape and trends, but simulations consistently predict

maximum coupling efficiencies which are approximately 3dB less than measured coupling efficiencies.

If the measurement and simulation coupling efficiency plots are normalized so that maximum coupled

powers are equalized, the axial separation and lateral misalignment curves generally agree to within 15%.

For small axial separations, this agreement falls to within 50% for some lateral misalignment cases. This

deviation may be associated with additional source mode coupling into receiving fiber modes as the

remaining unguided portion of the source mode field distribution diffracts within unguided fiber modes.

This effect would also explain the absolute difference in coupled power found between measurement and

simulation. Even with these deviations, the single-mode to multimode coupling case demonstrates a

prediction of fiber coupling efficiency based solely on physical fiber parameters and the coupled light

wavelength.

This modeling approach may be extended to fibers of arbitrary refractive-index profiles which satisfy the

weakly-guided approximation, and rectangular waveguides. Phase coherence effects can be modeled for

short fibers, and lenses and other elements with well-defined ABCD matrices can be modeled. Finally,

this method is computationally efficient, especially when considering the large number of modes involved

in the coupling process.

4. Conclusion

We have developed a modal coupling approach for modeling optical coupling between misaligned optical

fibers. Simulated coupling efficiencies for axial separations and lateral misalignments and give good

agreement with measurement results, although second-order effects limit accuracy at short distances. The

model predicted power-normalized coupling efficiencies for axial separations and lateral misalignments to within 15% for large axial separations. The approach is also applicable to laser diode-to-multimode fiber

coupling and fiber-to-waveguide coupling.

ACKNOWLEDGMENTS

This work was supported by ARPA Optoelectronics Technology Center (contract #MDA972-95- 1-0002)

and Cornell Industry Electronic Packaging Alliance.

REFERENCES

[1] S. C. Mettler, "A general characterization of splice loss for multimode optical fibers," Bell System

Technical Journal, vol. 58, no. 10, pp. 2163-2182,

Oct. 1979.

[2] P. Di Vita, and U. Rossi, "Realistic evaluation of

coupling loss between different optical fibers,"

Journal of Optical Communications, vol. 1, no. 1, pp. 26-32, Jan. 1980.

[3] Y. Daido, E. Miyauchi, and T. Iwama,

"Measuring fiber connection loss using steady

state power distribution: a method," Applied

Optics, vol. 20, no. 3, pp. 451-456, Mar. 1981.

[4] J. Sutherland, G. George, and J. P. Krusius,

"Optical coupling and alignment tolerances in optoelectronic array interface assemblies," Proc. of the 45th Elec. Comp. & Tech. Conf. , pp 577-

583, 1995.

[5] J. Sutherland, G. George, S. Van der Groen and J. P. Krusius, "Alignment tolerance measurements

and optical coupling modeling for optoelectronic

array interface assemblies," Proc. of the 46th

Elec. Comp. & Tech. Conf, pp 480-486, 1996.

[6] M. S. Cohen, G. W. Johnson et al, "Low-cost

fabrication of optical subassemblies," Proc. of the 46th Elec. Comp. & Tech. Conf, pp 1093-1100,

1996.

[7] J. M. Trewhella, M. M. Oprysko et al, "Unibody

plastic injection-molded optical sub-assembly for large core fiber," Proc. of the 46th Elec. Comp. &

Tech. Conf, pp 1116-1121, 1996.

[8] G. Keiser, Optical Fiber Communication,

McGraw-Hill, New York, NY, pp. 123-124, 1983.

92
Dispersion-compensating fibers witli high birefringence

J. Vobian, G. Herchenroder (FTZ Damistadt), K. Mori (IPHT Jena)

Abstract

For different applications, dispersion-compensating fibers (DC fibers) with high birefringence

may be very usefUl and advantageous, for instance, in fiber lasers in ring laser, but also in Fabry-Perot configurations Polarization-Maintaining Dispersion-Compensating (PM-DC)

fibers can used for the compensation for chromatic dispersion and stabilizing of the output

pulses by defined polarisation controlling. Also for several sensor applications PM-DC-fibers

can be installed with great benefit. In this paper, we present two PM-DC fibers with very high dispersion values in the third

optical window and, additionally, with very high birefringence values due to the elliptical cores (shape-induced or geometrical birefringence).

Tailoring of the PM-DC-fibers

We have presented an iterative, inverse method for calculating the refractive-index profile and

radius parameters of fibers with respect to prescribed dispersion and mode-field diameter

values (Petermann II) in the vicinity of 1550nm making use of the Levenberg-Marquardt

algorithm on the basis of the scalar Helmholtz equation and the singular Sturm-Liouville

eigenvalue solution [1], [2].

REFRACTIVE INDEX PROFILE

RETRnCTIVe INDEX PROFILE

" — 1 as 1 .,t

.03

.a?

M

r

radiu* t\jm1

Fig. 1 n(r)-profiles of the step-index fiber and the W-fiber

The high birefringence was obtained by elliptical cores of

the fibers taking advantage of shape-induced or geometrical

birefringence instead of stress-induced birefringence of

commercial Polarization-Maintaining (PM) fibers. The PMDC fibers (79M and 82M) were fabricated by MCVD at the

Institute of High Technology, Jena, Germany [1], [2]. Fig. J shows the two fiber profiles. As a consequence of the small core diameter, it is not possible to measure a quantitatively correct n(r)-profiIe using the convential Refracted Near Field

(RNF) technique. The core level n(r)*0 cannot be resolved

and is, therefore, too low. For the n(r) measurements, a part

of the preform was drawn to a fiber with a cladding diameter

ofSOO^im. Fig. 2 shows the fiber geometry of the W-fiber

82M.

Fig. 2 Core geometry of the

W-fiber (elliptic fit)

93
Measurements of the chromatic dispersion D(X,)

The measurements were performed making use of the Raman-fiber laser (RFL) and a MachZehnder interferometer (MZI) [3]. As a consequence of the necessary fit-function re- presentation, an extrapolation beyond the measuring range is forbidden. For this reason and for

the improvement of the accuracy in the third optical window and the reduction of the influence

of the fit fianction, it is important to establish a high upper limit of the wavelength range and to have as many measuring points as possible. In the case of the RFL this upper limit is 1740nm,

the maximum number of measuring points amounts to 30. Other advantages of this set-up are a good reproducibility and a high accuracy, the standard deviation of the measured group delay

data is comparable with that of interferometric methods. No spectral gap exists as in the case of semiconductor- or LED-set-ups. D(^) measurements of DC fibers with high negative

dispersion values are easily possible.

Measurements on short fiber pieces can be performed with the aid of a Mach-Zehnder

interferometer (MZI) with an incoherent radiation source [3]. The upper limit of the measuring

range is high (1700nm), a large number of measuring points can be obtained (up to 30) which

guarantees a high accuracy, again no ^-gap exists. Dispersion measurements on DC fibers with

L ps/kiTi*mi'i

0 1

1

1 L 1

1 f_

1100 1200 1300 1400 1500 1G00 1700

- wavelength in mn

-50 -

-100 -

-150 -

-200 -

-250 -

-300 LFig. 3 Chromatic dispersion of the W-fiber, Raman-fiber laser measurement

very high negative dispersion values are possible, but not trivial. Because of the large

difference of the group delay in both arms of the interferometer, normally, no interference

patterns can be observed. Two ways out are possible: shortening of the fiber under test (in

general, <2cm length difference between reference and test fibers) and/or lengthening of the

optical path in the air gap in the reference arm by an optically dense medium (Schott SF59

glass cylinder with n»2).

In Table 1 the dispersion measurements are compiled. The differences between RFL- and

MZI-results can be commented by fiber inhomogeneities, the measurements were performed on long fibers and short fiber pieces, respectively. But, in general, the agreement is satisfying.

Fig. 3 shows the dispersion curve of the W-fiber measured with the RFL. Important are the

very high negative dispersion values in the third optical window and the negative slope of the curve [2].

As mentioned above, both fibers have also a high birefringence. As a consequence, the RFLlaser pulse launched into the fiber is split into two pulses {Fig. -/), >^=1550nm, fiber length L is 6I5m.

94
PM-DC fiber Measuring

sct-up D(1300nm) m ps/km*nm

D(I540nm)

in ps/ktn*nm

D(I550nm)

in ps/km*nm

Hlliplic fit large/small axis Raman-fiber

laser -150.7 -194.7 -197.8 /.48/1.54nm

Mach-Zchnder

IIIILI IClUIIlLlCr

-148.3 -192.2 -195.3

fiber 79M

Raman-fiber

laser

-160.5 -105.0 -102.2

0 0/1/1 TAiitTi Mach-Zehnder

interferometer

-154.8 -101,2 -98.8

Table 1 Chromatic dispersion measuring results for two PM-DC fibers 79M and 82M

Fig. 4 Double-pulse presentation, step-index fiber,

>.=1550nm,

fiberlength L=615m

a sea leee isaa ieee asea aaea 3502 *eee 450B 5000

• IB— 12 9

The distance of the two pulses (1619ps) is a measure of the polarization-mode dispersion

(PMD) which is 2.63ps/m in this case. PMD measurements with a Michelson interferometer

The PMD measurements with the Michelson interferometer are performed in the time domain

[4] with an incoherent LED light source (AA^60nm) [5],

The relations between the group birefringence Bq, group beatlength 1q and the PMD

(Axgp) are: Ar,,=^^X (1)

L c lo-^, (2)

In Table 2 the measuring results are compiled, the PMD-, Bq- and l^-values for 4

wavelengths. For comparison the results of measurements with the Michelson interferometer on commercial PM (High Birefringence) fibers are added (PANDA-, YORK bow-tie- and

HiBi-AT&T-fibers). These latter measurements are in very good agreement with results

obtained by earlier measurements making use of the MZI and the Wavelength-Scanning

Method [6]. It is obvious that the birefringence and PMD values of the W- and step-index

fibers are remarkably higher and the group beatlength clearly lower than those of the

commercial PM fibers with their stress-induced birefringence which depends only weakly on

the wavelength. On the other hand the Bq- or PMD-dependence of the W- and step-index

fibers is evident. For the 79M fiber, PMD and Bq decrease with increasing wavelength, in the case of fiber 82M both parameter."^ increase with X. These results demonstrate the complicated

wavelength dependence of the birefringent fibers.

95
Wavelength in nm 1 / /O 1 7? 1

I -jZ 1

1 ^ 1 A 1 J 1 01 ^Af> I J 40

Parameter Fiber PMD in

ps/m

W-fiber 82M 5.64 5.91 8.68 8.89

Step-index fiber 79M 3.34 3.28 2.67 2.56 PANDA fiber 1. 44 1 .47 YORK bow-tie fiber 2. 18 2.22 AT&T-PM fiber 0. 77 0.78

Group

birefringence Bg*10-'

W-fiber 82M 16.92 17.73 26.04 26.67

Step-index fiber 79M 10.02 9.84 8.01 7.68 PANDA fiber 4. 32 4.41) YORK bow-tie fiber 6. 54 D.OO AT&T-PM fiber 2. 30 I.J J

Group

beatlength

in mm

W-fiber 82M 0.75 0.75 0.58 0.58

Step-index fiber 79M 1.27 1.34 1.89 2.01 PANDA fiber 3. 01 3.42 YORK bow-tie fiber 1. 97 2.25 AT&T-PM fiber 5. 58 6.46

Table 2 PMD, group birefringence and group beatlength of the two PM-DC fibers and

three commercial PM fibers for comparison

It is interesting to state the perfect agreement of the PMD value obtained with the double

pulse approach with the corresponding value in Table 2.

References

[1] R. Boness, J. Vobian, S. linger, J. Kirchhof Tailoring of Dispersion-Compensation

Fibers with High Compensation Ratios up to 30. Technical Digest Symposium on

Optical Fiber Measurements, Boulder 1994, p. 181-184

[2] R. Boness, W. Nowak, J. Vobian, S. linger, J. Kirchhof Tailoring of dispersioncompensation fibres with high compensation ratios up to 30. Pure Appl. Opt. 5 (1995), p. 333-344

[3] J. Vobian: Chromatic and Polarization Dispersion Measurements of Single-Mode

Fibers with a Mach-Zehnder Interferometer between 1200 and IVOOnm. J. Opt.

Commun. 11 (1990) 1, p. 29-36

[4] N. Gisin, J.-P. Von der Weid, J. -P. Pellaux: Polarization Mode Dispersion of Short

and Long Single-Mode Fibers. J. Lightw. Technol. Vol. 9, No. 7 (1991), p. 821-827

[5] V. Reichel, J. Vobian, N. Gisin; Polarization mode dispersion measurements on

long buried cables. Proc. EFOC&N 93, The Hague 1993, p. 56-61

[6] Ch. Chojetzki, H. Schmitzer, J. Vobian, W. Dultz: Four interferometric methods

for the determination of the birefringence and polarization dispersion of polarizaation maintaining optical fibers. Technical Digest Symposium on Optical Fiber

Measurements, Boulder 1992, p. 155-158. NIST Special Publication 839

96
Accurate wavelength calibration for optical spectrum analysers

D A Humphreys

National Physical Laboratory, Teddington, Middlesex, TWl 1 OLW, UK

Introduction

An Optical Spectrum Analyser (OSA) is a versatile instrument for measuring optical power as a function of

wavelength (and polarization in some instruments). Because of their widespread industrial use, the diversity

of applications and the mechanical nature of the design there is considerable interest in OSA calibration. To

meet this industrial need, we are developing calibration techniques for these instruments and participating

in the International Electrotechnical Commission (lEC) working group to standardise procedures for OSA

calibration. The key parameters of interest are:

Centre wavelength accuracy

Linearity of wavelength scale Instrument response

Resolution bandwidth

Power response

Power linearity (with power and wavelength)

Pulse performance

Polarization sensitivity

The wavelength accuracy and linearity of the wavelength scale are important parameters for users, particularly so for the development of wavelength division multiplex systems where the separation of the optical sources may be 1 nm or less. Typically the accuracy of the OSA will be limited by mechanical

constraints. These may be corrected at the design stage using software algorithms and a calibration procedure.

The wavelength scale of an OSA can be calibrated using a variety of fixed wavelength sources.

Spot wavelength calibration

There has been considerable interest in realising suitable reference standard lines at the 1.5 |im optical fibre

band'-^. These include gas lasers and atomic opto-galvanic transitions and molecular absorption features^ The

width of these lines is typically 500 MHz - 1 GHz so a single frequency tuneable laser locked system will have an overall accuracy of the order of tens of MHz.

At NPL, we use an opto-galvanic transition in Krypton'' at 1547.825 nm (DFB laser) and a CO absorption at 1560.5 nm (external-cavity laser)^. The accuracy of these systems significantly exceeds the requirements for OSA calibration.

Emission lines from discharge lamps give a series of line spectra and provide a lower cost alternative calibration method. The intensity of these lines is low, typically 1-20 pW depending on the line and optical

fibre type, but is sufficient to provide a calibration. A list of higher intensity lines^-^ for noble gases is shown

in table 1. Lines falling in the optical fibre communications wavelength range have been highlighted and lines used in this work are printed in bold type.

Wavelength scale linearity

Two approaches have been considered for the measurement of wavelength scale linearity: A tuneable laser used in conjunction with a wavemeter or calibrated (or self-calibrating) Fabry-Perot Etalon (FP) transmission

artifact. The advantage of the calibration artifact is that the response can be modelled and results calculated

for a wider range of wavelengths. The tuneable laser/wavemeter is useful for applications where the OSA is used over a restricted wavelength range. The calibration can be performed prior to use to give the highest

97
level of accuracy.

Table 1. Selected emission lines for noble gasses

Atom Wavelength Atom Wavelength Atom Wavelength Atom Wavelength

Kr 810.659 Ne 1029.824 Ne 1198.819 Xe 1414.631

Kr 811.513 Ar 1047.292 Ne 1206.964 Xe 1424.485

Kr 819.231 Ne Ne 1 246.28 Kr 1443.074

Kr

Kr

826.551

830.039

Ne

He

1080.103

1083.322

Xe

Ne

1262.684

1269.267

XT' Xe

Kr

1473.683

1473.846

Kr 851.121 He 1083.331 Ne 1291.555 Kr 1476.671

Kr 877.916 Ne 1084.745 ifill illlllm Ne 1523.488

Kr

At

893.114

912.547

Ne

Ne

1114.607

1115.081

Ne

At

1322.286

1331.685

Kr

Kr

1524.378

IS33.9I5

Ar

Ar

922.703

935.679

Ne

Ne

1139.355

1141.226

Ar

Ar

1337.077

1350.788

Kr

Xe

1537.624

1542.261

Ar

Ne

966.043

966.807

Ne

Ne

1152.59

1152.818

Kr

Kr

1362.614 Xe

Xe

1605.767

1673.272

Kr 975.443 Ne 1153.95 Xe 1366.079 Kr 1678.971

Ar

Xe

978.719

980.239

Ne

Ne

1161.726

1177.001

Kr

Ar

1366.213

1372.233

Kr

Kr

1685.809

1690.137

Xe

Kr

992.591

1022.426

Ne

Kr

1179.227

1182.261

Kr

Kr

1374.261

1404.95

Kr

Xe

1694.043

1733.05

Resolution bandwidth

The resolution bandwidth, AA,(A,), is an important parameter for optical noise-figure measurements. Because

an OSA is a grating instrument, the resolution is limited by the size of the input and output apertures. As these

have a finite width, the light incident on the grating will not be completely parallel. If the input aperture

contribution is ignored, then the resolution of the instrument is proportional to the width of the aperture in front of the photodetector. At normal incidence the instrument bandwidth will be given by equation 1

// 2 where AA,, is the full width at half maximum of

2

^ ' the instrument function and 60 is the acceptance angle of the output aperture and m

is the grating order.

In practice the response will be approximately rectangular for wide resolutions and triangular/gaussian for the minimum resolutions. As expected, multi-mode and single-mode fibres have different instrument

AX = 266

^

98
responses.

The minimum resolution bandwidth of a commercial OSA is typically 0. 1 nm which corresponds to a frequency

of 12.5 GHz at 1550 nm.

Measurements

An OSA was calibrated over the range 1 180 nm - 1680 nm using six lines from a high-intensity Kr discharge

lamp. The linearity over the intervening points was measured using a calibration artefact and two LED sources. The measured results (figure 1) were normalised (figure 2) for source intensity variation and system sensitivity

using equation 2: where R„„^ is the normalised artefact response and the optical powers are in R

P - P artefact noise Preference - Pnoise Watts..

The onission lines were used to provide calibration for both the OSA and the artefact. The calculations were

performed in three stages:

1. Estimate the FP spacing from measurements

around each of the Kr emission lines.

2. Calculate the true value of the spacing from

measurements at the Kr emission lines.

artefact response ^

system noise ^

1)00 1400

3. Determine the wavelength scale linearity.

The calculated linearity with wavelength (figure 3) give the correction factor for air and vacuum wavelengths.

Figure 1 Source-power, artefact response and system

noise results

The variation of the resolution bandwidth was also investigated as a fiinction of wavelength. A rectangular fit was used for the 1 nm banwidth and a trianglular fit was used for the 0. 1 nm bandwidth. The results (figure 4)

indicate that at 0. 1 nm the wavelength dependent effects are well balanced but at 1 nm the grating angle effect

is dominant.

* 0.9

198 204

Frequency, THz Figure 2 Normalised artefact response over limited range

99
Summary

The work at NPL to provide traceable calibration for OSA wavelength scales has been outlined. Calibration

results over the range 1 180 - 1680 nm for a commercial OSA have been presented.

1

f

-0.4

-0.1

-0.8

,:„:

1

a

0

a

1100 1300 1300 1400 ISOO

Figure 3 Wavelength correction factor air (line),

vacuimi (points)

ISOO 1700

I

0.6

0

D —

9

0

0

0

0

0

7/00 1100 1300 1400 1500 1600 1700

woifdength, m

Figure 4 Wavelength dependence of the resolution

bandwidth

References

1. D J E Knight et al.: 'Frequency standards for the optical-fiber communication bands: a review of

user requirements and of physical possibilities for meeting them in practice', OSA 1993 Annual

Meeting, 3-8 October 1993, Toronto Canada. Paper MBl (invited), pp. 2-3. A I Lucero et al: 'Survey of atomic transitions for absolute frequency locking of lasers for lightwave systems', IEEE Photonics Technology Letters, Vol. 3 (5), pp. 484-486, May 1991.

F Bertinetto et al: Frequency stabilization of DFB laser diodes to Hydrogen, Iodide at 1.5364 \im\

OFC 1993.

4. D J E Knight et al: 'Frequency measurements on HI and Kr line at 1.548 |im', 10th European

Frequency and Time Forum, 5-7 March 1996, Brighton, UK

5. D J E Knight et al: 'Absolute frequency measurement of the 3-0 R(2 1

) transition of CO at 1 .5605

\im for optical communication standards', Proc. EFTF 95, Besangon, March 1995.

6. CRC Handbook of Chemistry and Physics, 70th edition, R C Weast and D R Lide (ed), CRC

Press, 1989.

Elden: J. Opt. Soc. Am., 1953, 43, 339

100
Nonlinearity of optical fiber power meters

Igor Vayshenker, Shao Yang*, Xiaoyu Li, Thomas R. Scott

National Institute of Standards and Technology, 325 Broadway, Boulder, CO 80303

Ohmeda Medical Systems, Inc., Louisville, CO 80027

Abstract

We have developed a system for measuring the nonlinearity of optical fiber power meters over a dynamic range of more than 60 dB at three telecommunications wavelengths. This system

uses optical fiber components and is designed to accommodate common optical powers; it is based on the triplet superposition method. This system provides accurate determination of optical power

meter nonlinearity through the use of correction factors.

1. Introduction

The basic assumption for accurate optical power measurement is that the meter output

reading is directly proportional to the optical input power. This proportionality property is called

linearity, and the departure fi^om this direct proportionality is defined as nonlinearity. Optical power

meter nonlinearity is defined as the relative difference between the responsivity at an arbitrary power

and the responsivity at the calibration power:

^

R(P) - R(P ) ^nl(P;P ) =

^

'

' (1)

where R(P)=V/P is the responsivity of the meter at optical power P, the subscript c indicates the

calibration point, and V is the meter output. A fianction that describes the relationship between the

incident optical power P and the optical meter output is called the response fianction. The inverse

of the response fianction is called the conversion fianction and is depicted in Figure 1

Figure 1 . The conversion fianction.

101
It is often more convenient to use the conversion function rather than the response function.

We can define the nonlinearity in terms of the output V as P(V )V

When the nonlinearity is small, a polynomial can represent the conversion function sufficiently well.

While the calibration gives the true input power from the power meter reading (output) at the calibration point, the measurement of nonlinearity and range discontinuity, together with

calibration, provides this input-output relationship at any power within the entire dynamic range of

the power meter.

It is, therefore, convenient to express the measured nonlinearity in terms of the conversion

function P=P(V), which relates the input power P to the output V, and referenced to the calibration

output V^.

1.1. Correction Factor for Nonlinearity and Range Discontinuity

The true input power P is obtained from the power meter reading V by

(3)

where ¥=Y^/P^ is the calibration factor.

a,[c] ^^g'^t^l^'

CF=-1—X

^

k-1

a, [ml " ^ l-EbJm]V'^-^ (4)

k=2 A^[i.V(V;0)Al(Vc;0)],

a correction factor due to nonlinearity and range discontinuity, m is a number that corresponds to a specific range of an optical power meter, a^ and b^ are coefficients, and c corresponds to the

calibration point. Consequently, each range of a power meter has its own correction factor. The

degree of polynomial, n is usually equal to 3 or 4 depending on the size of the data standard

deviation.

102
2. Measurement system and results

We have based the operation of our system on the triplet superposition method which

relies on the principle that, for a linear meter, the sum of meter outputs corresponding to inputs from

two individual beams should equal the output when the two beams are combined and incident on

the meter at the same time. The measurement system is depicted in Figure 2. We use a high-power,

single-mode, fiber-pigtailed diode laser whose output power is stabilized; the laser is temperaturecontrolled. An external optical attenuator with a dynamic range of 60 dB is used to provide variable

optical power. The optical power from the attenuator is divided into two approximately equal parts

by introducing a 3 dB fiber splitter; one of the splitter arms has an additional length of fiber to avoid

interference. A computer-controlled shutter is inserted into a collimated-beam section in each arm. Both signals are recombined in a 3 dB fiber coupler. We use single-mode fiber components

throughout the system. We call a platform with fiber components and two shutters a switching

matrix. The measurements were performed by taking sets of three power readings from the test meter: (1) shutter 1 is open and shutter 2 is closed, (2) shutter 1 is closed and shutter 2 is open, and

(3) both shutters are open. This sequence is then repeated at different powers.

Figure 2. Measurement system.

We constructed two switching matrices: one for 850 nm, and the other for both 1300 nm and

1550 nm. Both switching matrices shared one attenuator. To measure the range discontinuity

(offsets between range or scale settings), readings are taken at the lower power end of each range

and compared to the readings on the higher power region of the next lower range (if any) at a constant power. The calculated correction factors result from the meter nonlinearity within each

range, combined with the range discontinuity.

Figures 3 and 4 depict correction factors obtained on a typical optical power meter at 850

nm and 1300 nm, respectively. Each data group represents a separate power range of the meter. The

same optical power meter was significantly nonlinear at 850 nm (Figure 3) and very linear at 1300

103
nm (Figure 4)

^ O 99 " o.se

O.OOIS 0.01S 0.1S 1.5 15 ISO 1 200

Flower (pW)

Figure 3. Correction factor vs. power at 850 nm.

1 .0002

1 ^ 0.9996

o 0.9992

0.999

0.99SS0.0004 0.005 O.I 25 2.5 3S 500

F=»ower (nW)

Figure 4. Correction factor vs. power at 1300 nm.

Acknowledgments

This work is supported by the Navy's Metrology R&D Calibration Standards Program

under the management of Joe Dressel (NAVSEA 043 12).

References

1. "Calibration of Fiber-Optic Power Meters," Draft International Standard, EEC TC 86, June

25,1992.

2. S. Yang, I. Vayshenker, X. Li, and T.R. Scott, "Optical detector nonlinearity: a comparison of

five methods," Digest, Conf Precision Electromagnetic Measurements, pp. 455-456, June-July

1994.

3. S. Yang, I. Vayshenker, X. Li, M. Zander and T.R. Scott, "Optical detector nonlinearity:

Simulation," Natl. Inst. Stand. Technol. Technical Note 1376, May 19, 1995.

4. I. Vayshenker, S. Yang, X. Li, and T.R. Scott, "Automated measurement of nonlinearity of

optical fiber power meters," Proc. SPIE, Vol. 2550, pp. 12-19, San Diego, CA, July 1 1-12, 1995.

104
Methods of Suppressing Stimulated Brillouin Scattering in Optical Fibers by Manipulation of the

Fiber Properties

Clifford Headley, J. B. Clayton, and W. A. Reed

I INTRODUCTION

Stimulated Brillouin scattering (SBS) is a nonlinear optical process. Its occurance in optical

fibers is characterised by an energy transfer from an input (pump) wave to a backward-traveling

frequency-downshifted (Stokes) wave. [1,2] This occurs because the pump light generates an acoustic wave through the process of electrostriction. The acoustic wave causes a periodic modulation of the fibers refractive index creating a Bragg grating from which photons can be scattered. Because both energy and

momentum must be conserved, counterpropagating photons are preferentially amplified. The frequency of

the photons that are preferentially amplified is slightly less than that of the pump wave, due to the Doppler shift associated with the grating moving at the acoustic velocity. If a wave with the appropriate

frequency shift is propagating coimter to the pump wave it will be amplified. However, even in the absence of an input signal, SBS generates a Stokes wave from spontaneously scattered photons.

Both active and passive methods have been anployed to suppress SBS. Active methods involve

broadening the spectral width of the input pulse, [2] while passive methods involve modulating the fiber characteristics. [3 -20] It is the passive techniques \\liich are discussed here. Section n defines several terms and parameters necessary to understand the approaches described. Section UI reviews the suppression techniques, and Section IV provides an experimental example. Fmally, Section V

summarizes.

n BACKGROUND

Stimulated Brillouin scattering is characterized by several parameters. These include the

Brillouin frequency shift, Vb, given as [1,2]

2nvA ... Vb=—, (1)

where n is the refractive index for light at the pump wavelength, Va is the velocity of the acoustic wave in the optical fiber, and \ is the wavelength of the pump light. Vb represents the frequency difference between the pump and Stokes wave. The Brillouin frequency shift is typically on the order of 10 GHz at a wavelength of 1.55 iim in silica. The growth in the energy of the Stokes wave at frequency v is characterized by the Brillouin-gain

coefficient gB(v). Assiiming that the acoustic waves decay at a exponential rate characterized by the time

coefficient Tb, and that the pump laser is operating in the continuous wave (cw) or quasi-cw (pulse width

To » Tb) regime the Brillouin gain spectrum has a Lorentzian profile [1], with a peak value at v = Vb. The Brillouin gain spectrum is also characterized by its full width at half maximum, Avb, which is related

to the lifetime of the acoustic waves by the expression Avb = (nT b)~^ Avb is typically tens of MHz, and

this relatively narrow spectral linewidth aids suppression techniques. If the assumption of the cw or

qviasi-cw operating regime is removed (pulse width To < Tb) studies show that the SBS gain is significantly reduced, so that other nonlinear effects become important. [3] In order to compare the effect of SBS between different fibers a figure of merit is necessary. Such

a measure can be somewhat arbitrarily chosen. However, a convenient and often used condition is a threshold input power, P^, defined as the input power necessary for the Stokes light at the input end of the

fiber to equal the transmitted pump light. Assuming a Lorentzian gain profile this criterion is given by,

II]

105
gB(VB)P L 7-^= 21k, (2)

where A,ff is the efifective core area of the fiber, and Leff is the efifective interaction length of the fiber and is given by [1], L^g = a"'[l-exp(-aL)] where a is the loss coefficient of the fiber and L the actual fiber length. The factor k in Eq. (4) has a value between 1-2 depending on how well the pump and Stokes wave

maintam their polarization states along the fiber. A value of k = 1.5 occurs when the polarization is completely scrambled. [21] For input pump powers greater than Pth the effects of SBS are considered

significant.

Pfli is a good figure of merit because it allows the calculation of the maximum power one can put

mto a fiber before SBS becomes significant. However, it has several disadvantages. The first shortcoming

is that Pa, mcreases with loss giving the impression one has a better fiber. To overcome this difficulty, a

criterion which may be used is to specify an output power, = Pu,(L) - oL, which is the amount of power

that would be transmitted through a fiber of length L when Pth is the input power. The second drawi^ack

is that Pa, is length dependent, and therefore does not readily allow the comparison of fibers of different lengths. To overcome this difficulty an increase in power Fj^c may be specified. Pi„, represents the

increase in pump power mto a fiber prepared using an SBS suppression technique compared to the same

length of fiber prepared without applying SBS suppression. The inaease in SBS threshold is then valid

for arbitrary lengths of the fiber. A final note about comparing the effects of SBS fi"om one fiber to a next. gBiya) is dependent on

the spectral width of the pump source, a fact that is used in active SBS suppression techniques. Hence the

difference in pump spectral width must be accounted for when comparing fibers tested with different pump sources. m SUPPRESSION TECHNIQUES

With the definitions given above an imderstanding of the passive methods used to suppress SBS

can be obtained. These techniques rely on varying either the waveguide or material properties of the fiber so that the Brillouin fi-equency changes before the Stokes light at a given fi-equency reaches threshold. [4-

19]. Because of the narrow linewidth of the Brillouin-gain curve, a small change in Vb means that a new

Stokes fi"equency is being amplified, thereby reducing pump depletion. Equation (1) shov^ vb can be

changed by varying n or Va. The fiber properties can be modulated both during (fabrication processes) or after (external

modulation) fabrication. As a general principal, fabrication processes have the advantage that the SBS

suppression is an inherent property of the fiber and no additional effort is needed during fiber installation. However, the downside of fabrication processes compared to external modulation is that they tend to have

a more drastic effect on other fiber properties such as loss or dispersion. External modulation techniques can be more easily applied to fibers such as dispersion compensating fibers (DCF) v/hich are sitting in a

hut and more accessible and also have less fi-eedom in the design parameters. In what follows several techniques will be discussed. These methods include varying the fiber core area,[4-6], fiber-dopant

concentration,[7-l 1] strain along the fiber,[12-16], inserting optical isolators along the fiber,[16] and

temperature of the fiber[ 18-20],

III. 1 Fabrication Processes

During the fabrication process there are several properties of the fiber v^diich can be modulated.

These include the fiber core radius, fiber dopant concentration and residual strain induced on the fiber by

the drawing process.

III. 1 .a Core Radius

Varying the fiber core radius along the fiber leads to a reduction in SBS. This results because a

given fiber supports certain acoustic modes, [22-25] which travel at acoustic velocities determined by fiber

106
parameters such as core area. Changing Va leads to a varying vb and hence a reduction in SBS. It has

been shown that vb varies inversely with the core diameter. [4-6] In an experiment performed by Shiraki

et. al. at a pump wavelength of 1.55 )im, a 14 km length of fiber whose mode field diameter changed from

8.1 1 |im to 6.93 |im showed a change m Brillouin frequency shift, 5vb, of 49 MHz.[4] This resulted in an

increase in Pth of about 3.6 dB compared to similar fibers with constant radii. It should be noted that the appearance of multiple peaks in the SBS spectrum is due to more than one acoustic mode contributing to the SBS process. From a practical standpoint, one problem in the fabrication of this fiber is that changing the core

radius ofthe fiber affects other fiber properties such as dispersion and loss. The dispersion of an ordinary

step index fib«- increases with increasing core area. [5] In addition, the fiber diameter used in systems is a

standardized value, and achieving a low-loss splice between two fibers of dissimilar areas is very difficult.

in. 1 .b Dopant Concentration

By changing the dopant concentration along the length of an optical fiber SBS effects can be

reduced. [7-1 1] Both Va and n are affected by the dopant concentration, but the majority of the shift in Vb

arises form variation in Va. This is because the material density, p, of the fiber is related to the amoimt of

dopants present, and the velocity of sound in turn depends on p. [7] As afready indicated a varying

acoustic velocity modulates vb and leads to the suppression of SBS.

In an experiment performed by Tateda et. al. [9] and by Shiraki et. al. [10, 11] with fluorine as a dopant the refractive indices of the core and the inner cladding were monotonically lowered keeping

the relative index diffwaice between the core and cladding the same. While the exact variation will depend on fiber type, in [10] a 7 dB increase in Pa, was observed for a 350 MHz shift in Vb. By taking the

derivative of Eq. (1) with respect to dopant concentration, and using the values of 356 MHz and 0.2 % for the changes in Brillouin frequency and fiber refractive index respectively, this calculation shows that it must be the change in acoustic velocity that leads to a change in vb. By fabricating fibers with various

dopant concentrations,[ll] vb was shown empirically to vary as Vb = 11.045 -0.277(F) + 0.045(GeO2)

[GHz].

One of the major drawbacks of this technique is that like changing the core radius, changing the dopant concentration affects other fiber properties. When F is added to a silica fiber the refractive index

of the fiber decreases whereas Ge02 increases n. As the index of the fiber is increased along the fiber length the fiber loss also increases. The opposite relationship between n and the dispersion holds true. [8]

m.l.c Strain By varying the strain induced on a fiber during the draw process, the Young's modulus (defined

as the ratio of stress to strain) of the fiber is modulated. The acoustic velocity of the fiber depends on the Yoimg's modulus, therefore, changing the strain along the fiber raises the Brillouin threshold. A periodic

residual strain can be applied to the fiber by varying the draw tension on the fiber during

fabrication. [12,25] In an experiment performed by Wada et. al. the drawing tension was varied rectangulary with an

amplitude of 16 g every 2.5 km for a 23 km length of fiber.[12] A predicted 75 MHz change in Vb led to a

3 dB increase in the Brillouin threshold power. In this experiment, a fluorine-doped cladding\silica core

fiber was used because the higher viscosity of the core caused it to solidify first during the drawing

process, so that most of the residual strain due to draw tension will be in the core. The reverse situation is true for the case of germanium-doped core\silica clad fibers, and the

effect of strain in this type of fiber is reduced. This is the major drawback of this technique. [12] A

second consideration is that there is a possible reduction in the lifetime of the fiber associated with

introducing a higher strain in the fiber.

IIL 2 External Modulation

Three external modulation techniques will be discussed here, varying the fiber strain, inserting

isolators along the fiber, and introducing a temperature variation along the fiber.

107
m.2.a Strain The effect of strain on reducing the SBS threshold was described in Section UI.l.c, the difference here is that the strain is induced on the fiber after fabrication. [13- 16] The relationship between the

residual strain change, 58, and the frequency shift, 5vb(s), normalized to the unstrained frequency shift, vb(0), was calculated and verified for a variety of cladding/ core dopant values for both fluorine and

germanium doped fibers by Kurashima et. a/.. [16] A linear variation was measured, and while the exact value of this parameter depends on which dopant is used and how much of it, the variation with these parameters was fairly small[13-16].

hi one very clever experiment by Yoshizawa et. al. [14,15] twelve strands of optical fibers were wrapped around a coated wire. Four such bundles of twelve optical fibers were themselves then wrapped

around a larger coated wire. The effect of this double wrapping was to vary the strain along the fiber sinusoidally, which in turn varied the velocity of sound in the fiber leading to a 7 dB increase in the SBS

threshold. Moreover the fiber loss was measured and found to be equal to that before the wrapping.

The major drawback to this particular method of SBS suppression is that the lifetime of the cable

is reduced. In the experiment described above, carbon coated fiber was used because of the increased

strain. In addition unplementing the strain variation can be difiBcult.

III.2.b Isolators One of the most simple and obvious ways of reducing SBS is to introduce a directionally

dependent loss into the fiber so that any backscattered light is reduced whilst the input light is passed.

This can be done by the insertion of optical isolators along the fiber. [17] Under ideal conditions where

there is no isolator insertion the increase in the SBS threshold should be N + 1 where N is the number of

isolators. In an experiment performed by Takustina et. al. an isolator was inserted into a fiber, it was

shown that the transmitted power level for SBS was raised by 1.2 dB. This deviation from the predicted

3 dB ideal value is due to insertion losses and less than total isolation. This technique suffers from the problems of very minimal improvement, significant increase in the fiber loss, increased cost associated with buying isolators, and splicing into the fiber. However, this result made a very significant point, and that is that the ideal placement of the isolator is towards the

begmning of the fiber span, a principle which is usefiil for all SBS suppression techniques

ni.2.c Temperature

The Brillouin frequency shift experienced by a fiber is temperature dependent. If a length

dependent temperature change is applied across the fiber length then the effect of SBS can be reduced. [18-

20] Mathematically, the temperature variation of vb can be found by taking the derivative with respect to temperature, T, of Eq. (1). Both va and n have a temperature dependence, but the dominant contribution

to this frequency shift comes from the change in refractive index, with the change in acoustic velocity

accounting for only about 7 % of the total contribution.[19] The linear variation of Vb with temperature was found by Kurashima et. al. to be 1.33 MHz/°C for a silica core\flourine-doped cladding fiber, and for germanium-doped core/silica cladding fiber 1.17 MHz/°C, both measured with a 1.32 }im

wavelength. [18] Exact values for a fiber will depend on the dopant concentration. For both fibers the

core cladding index difference was 0.3%. Interestingly enough it is found that the Brillouin gain

bandwidth also decreased linearly as temperature increased.[18]

In another experiment conducted by Imai et. al. half of a fiber spool was heated and hence a temperature gradient was introduced along the fiber. [20] As the temperature gradient along the fiber was

increased the SBS threshold was raised by 3.5 dB corresponding to a 35 °C temperature difference across the fiber.There are two drawbacks to this procedure. The first is that this is more difficult to implement

for fibers that are distributed along a span. This technique more readily lends itself to fibers that that remain on a spool such as dispersion compensating fiber. It maybe especially usefiil for fiber spools in devices which can be packaged to take advantage of the heat generated by other devices, as some form of

power supply is necessary in order to supply heat to or cool the fiber

108
IV EXPERIMENTAL RESULTS

Figure 1 shows a typical experimental setup for measuring the increase in Brillouin threshold in a fiber sample. The output of the pump laser passes through a variable attenuator, which controls the power entoing Ae test Gber. From the attenuator the light goes through port 1 of a tap which nominally

s«ids 99 % of the light into tiie test fiber tiirough port 2, \^ile 1 % of the light is measured at p<Bt 3 as a

mcKiita- of the input signal. One perc«it of any light that is backscattered in the fiber enters the tap

through pOTt 2, is drtected on the 4th arm of the tap, and this signal is used as a measure of the Stokes

light. The loss aCTOSs each port and the exact percentage of light transmitted at ports 3 and 4 are

Pump

Laser Variable

Attenuator

Backscattei Power

Monitor

Input Power

Monitor

Transmitted Power

Monitor

Fig. 1 Stimulated Brillouin scattering measurement setup.

measured by inserting known power values into ports 1 and 2 respectively and measuring the power across the other arms. The Brillouin frequency shift is measured using a HP71400C Lighhtwave System

Analyzer.F«- illustrative purposes the use of temperature to increase the Brillouin threshold is studied. Twelve kilraneta's of dispersicsi aanpensating fiber was woimd cmto three sq)arate spools, with the first and last secticms each 3 km long. This allows portions of the fiber to be independently cooled in LN2, and

the COTresponding SBS threshold for that situatirai to be measured. When the input and output ends of the

fiber wffl-e cooled there was a 3dB increase in the The reason for this is seen in Fig. 2, A\iiich is a plot

of the nwmalized spectrum of the beat signal between the pump and Stokes signal, (a) when the

- - -U ncootad Fib«r In/OutEnds Coolgd 1.0 -

9.40 9.4S 9.60 9. 86 9.60 9.86 9.70

Fraqaancy (M HO

Figure 2. Plot of Brillouin frequency shift for (a) uncooled fiber (dotted) and (b) input and output ends

cooled (solid).

fiber is vmcooled and (b) when the input and output ends of the fiber are cooled These results show that cooling the fiber has the effect of shifting a portion of the Stokes energy by 150 MHz. This result corresponds to a tempo-ature coefficient of Brillouin frequency shift dvs/dT of 0.75 MHz/°C, a figure

109
lower than that given m Section ni.2.c due to the larger amount of Ge doping in dispersion compensating

fiber. [18] The shift in frequency of the Stokes energy leads to the reduction in SBS threshold.

V. CONCLUSIONS

Several techniques for reducing SBS both during and after fiber fabrication were reviewed.

These techniques included modulating the fiber properties such as core radius, dopant concentration, fiber

strain both during and after production, temperature along the length of the fiber, and inserting isolators. The advantages and disadvantages of each particular technique were given. For heuristic purposes, an

experiment to measure the increase in Brillouin threshold due to temperature changes was presented.

REFERENCES

1. G. P. Agrawal, Nonlinear Fiber Optics^ 2nd ed. (Academic, San Diego, 1995).

2. D. Cotter, J. Opt. Commun., 4, 10-19 (1983).

3. A. R. Chraplyvy, ffiEE J. Lightwave Technol., 8, 1548-1557 (1990).

4. K. Shiraki, M. Ohashi and M. Tateda, Electron. Lett., 31, 668-669 (1995).

5. K. Shiraki, M. Ohashi and M. Tateda, Proc. 21st Conf on Opt. Comm. 325-328 (1995).

6. M. Ohashi, N. Shibata and K. Shiraki, Electron. Lett., 28, 900-902 (1995).

7. R. W. Tkach, A. R. Chraplyvy and R. M. Derosier, Electron. Lett., 22, 101 1-1013 (1986).

8. M. Ohashi and M. Tateda, J. Lightwave Technol., 11, 1941-1945 (1993).

9. M. Tateda, M. Ohashi and K. Shiraki, OFC/IOOC 1993 Technical Digest paperThJ4.

10. K. Shiraki, M. Ohashi, and M. Tateda, J. Lightwave Technol., 14, 549-554 (1996).

11. K. Shiraki, M. Ohashi, and M. Tateda, J. Lightwave Technol., 14, 50-57 (1996).

12. A. Wada, T. Nozawa, D. Tanaka, T. Sakai, and R. Yamauchi, Proc. European Conf on Opt.

Commun. 1991 MoBl, (1991).

13. T. Horiguchi, T. Kurashima and M. Tateda, ffiEE Photon. Techn. Lett., 1, 107-108 (1989).

14. N. Yoshizawa, T. Horiguchi and T. Kurashima, Electron. Lett., 27, 1100-1101 (1991).

15. N. Yoshizawa and T. Imai, J. Lightwave Technol., 11, 1518-1522 (1993).

16. T. Kurashima, T. Horiguchi, and M. Tateda, 7th hitemational Conf on Integrated Optics and Optical

Fiber Communication, 4, 21C4-2 70-71 (Kobe Japan 1989)

17. Y. Takushima and T. Okoshi, Electron. Lett., 28, 1155-1157 (1992).

18. T. Kurashima, T. Horiguchi and M. Tateda, ffiEE Photon. Techn. Lett., 2, 718-720 ) (1990).

19. D. Culverhouse, F. Farahi, C. N. Pannell, and D. A. Jackson, Electron. Lett. 25, 913-915 (1989).

20. Y. Imai and N. Shimada, ffiEE Photon. Techn. Lett., 5, 1335-1337 (1993).

21. C. K. Jen. A, Safaai-Jazi and G. Famell, ffiEE Trans. Ultrasonics, ferroelectrics and frequency

control, 33, 634-643 (1986).

22. N. Shibata, K. Okamoto and Y. Azuma, J. Opt. Soc. Am. B, 6, 1 167-1 174 (1989).

23. W. Henry, hit. J. Optoelec., 7, 453-478 (1992).

24. M. Ohashi, N. Shibata and K. Shiraki, Electron. Lett., 28, 900-902 (1992).

25. U. Paek and C. Kurkjian, J. Am. Ceramic Soc. 58, 330-335 (1975).

110
Local Analysis of Stimulated Brillouin Interaction

in Installed Fiber Optics Cables

Marc Nikles, Luc Thevenaz, Pascal Salina*, Philippe A. Robert

Swiss Federal Institute of Technology, Metrology Laboratory, CH-1015 Lausanne, Switzerland

* Swiss Telecom PTT, Osternnundigenstr. 93, CH-3029 Bern, Switzerland

Brillouin gain spectrum measurement along an optical fiber has recently gained a lot of interests owing

to its potentiality for strain monitoring in installed telecom cables [1,2]. The purpose of this paper is to show that this potentiality is being now effective, since field measurements of installed fiber optics

cables currently in operation are demonstrated. A portable instrument has been developed, based on an

original experimental configviration developed in our Institute [3] that is shortly described below.

The Brillouin Gain Spectrum (BGS) in silical optical fibers is downshifted with respect to the pump light

frequency by the BriUouin frequency shift vg which takes values from 11.5 to 13 GHz depending on the

refractive index profile at a pump wavelength near 1300 nm [4]. The value of vg turns out to be very

strain sensitive, shifting vg by about 600 MHz per percent of elongation [5]. This measurement requires

two Ught waves propagating in opposite directions throughout the fiber, since Stimulated Brillouin

Scattering (SBS) amplification is possible only in the backward direction. One light wave pumps the

medium and the other acts as a probe signal, experiencing amplification when it Ues within the

BriUouin gain spectral range.

Pump Laser ^ Coupler

95/5

Cleavedend

I ) Fiber under

test

1 Data DET

Acquisition

Trig t

Fig.l: Experimental setup used for distributed temperature measurement.

Distributed measurements can be performed using a modified OTDR technique. The experimental

configuration shown in Fig.l is based on the pump and probe technique and presents the following

advantages for field measurements; 1) a single fiber end is accessed, 2) a single laser source is used, 3) few

optical elements are required.

Ill
An ultra-wideband integrated LiNbOs intensity electrooptic modulator (EOM) is the key element of the

setup, since it is used for: 1) pulsing the CW laser light from a 150mW Nd:YAG laser to form the pump

signal, 2) the generation and frequency tuning of the probe signal. The frequency shift of the laser light

is achieved by simply applying a CW microwave signal on the EOM electrodes. This creates sidebands

in the laser spectrum, so that the first lower sideband lies in the BGS generated by the pump, when the

modulation frequency/;,! is close to the Brillouin frequency shift Vg. The DC bias voltage on the EOM

basically just determines the amount of transmitted amplitude of the fundamental frequency. It is set, so

that the carrier is totally suppressed and the remaining probe signal is uniquely made of the

modulation sidebands. The CW probe and the pvmnp pulses are both launched into the fiber under test

at the same fiber end. The pump pulse provides gain to the probe signal during its forward

propagation through the SBS process while the probe signal is amplified on the way back. Only a few

percent of reflection is necessary at the fiber far end, so that Fresnel reflection is actually sufficient. Since

the first upper sideband is not relevant for the measurement and even has a negative effect on the

contrast, it must be suppressed using an optical filter. The probe intensity is monitored as a function of

time and the delay between the launch of the pump pulse and the detection of the probe signal gives

the positional information.

This technique makes the scan of the probe optical frequency very simple and convenient by just

varying the microwave signal frequency. Since this technique uses a single laser and a modulator for

the generation of the pump and probe signals, it insures a perfect stability of their frequency difference.

Measurements of the distribution of the Brillouin frequency shift along a sensing fiber can therefore be

performed with an excellent resolution using this technique. The measured standard deviation in the

determination of the Brillouin shift frequency is 300 kHz. The best spatial resolution obtained so far is

15 m for a standard fiber, given by the minimum pump pulse width leaving a sufficient gain.

Local analysis of stimulated Brillouin interaction (LASBl) measurements were carried out on a 6.8 kmfiber optics cable link between the Swiss Telecom exchange stations of Versoix and Anieres across the

lake of Geneva, Switzerland. A part of the cable is designed for underwater applications and contains 2

tubes with 10 single mode optical fibers. Each fiber link is actually made of 5 segments of fibers spliced

together. The topographical arrangement is schematically shown in Fig. 2a, the successive splices being

indicated by SI through S4. Starting from the exchange station of Versoix, the conduit route can be

described as follow: the first section of cable (560m) is a standard unarmoured telecom fiber optics cable,

then 60m of the same cable (S1-S2) is used to join the underwater cable (armoured reinforced cable). The

first 460m of the underwater cable goes through the lake of Geneva shore line and approximately

3700m are laid on the lake bed. On the other waterside, 360m of conduit brings the cable to a manhole

where it is spliced (S3) to a 550m standard telecom cable. Finally 590m of another cable joins from S4 to

the exchange station of Anieres.

112
Fiber optics cable

rsi S2

2.84

2.82

2.80

2.78

2.76

2.74

2.72

2.70

2.68

Underwater section

S3 S4

JmJ'*' I Tp jiii '^iii i ^ '1. 1

- b.

2.84 h

2.82

2.80

2.78

2.76

2.74

2.72

2.70

2.68

Local

Strain

0 1000 2000 3000 4Q00 5000

Position along the cable, m

6000 70Q0

Fig.2 Measurement of the Brillouin frequency shift on a underwaterfiber optics cable. Topographical

representation of the cable route (a). Local Brillouin frequency shift measured with a 20m spatial resolution for 2 different fibers of the cable (b and c).

113
Several optical fibers of the cable have been tested. Fig. 2b and 2c shows the BriUouin frequency shift vg along two fibers measured with a 20m spatial resolution. The different sections of the cable are

clearly identified by their different vg. The vg value can differ from fiber to fiber depending on the

fiber parameter [3]. On the terrestrial route, the cable is installed in a conduit. The temperature being relatively constant in

the conduit, the changes in the Brillouin frequency shift are attributed to local strain. On the contrary

the underwater portion is just laid on the lake bed and is consequently subjected to temperature

variations. As a matter of fact the temperature gradient of the water is clearly observed on the two plots

of Fig.2. The temperature is measured to be actually 5°C colder (frequency shift of 6.45 MHz smaller) at

the deepest point (61m depth, positions 2000m to 3000m in the cable) than the water temperature

nearby the shores (positions 1090m and 5300m). At 2600m away from the shore (position 3690m) the

depth of the lake goes up to 31m and an 1°C increase is detected.

Concerning the strain experienced by the fibers, the first and last sections of the cable are installed

through a winding route and undergo many strong curvatures. Nevertheless the variations of the

measured Brillouin frequency shift (Fig.2b) is 12.9MHz and 11.6MHz for the first and last section

respectively, corresponding to 2.15*10"* and 1.9*10"* equivalent strain. The section between S3 and S4

presents a slight slope that corresponds to a 1.1*10"* strain. In Fig.2c the peak at position 2100m in the

underwater section is attributed to a local strain. Even though the magnitude of the perturbation is not

perfectly resolved because of the spatial resolution of the system (20m), the problem is easily detected.

However the local strain is probably higher than the measured 2.7*10'^. A portable instrument for local Brillouin gain spectrum analysis has been built and was used for on site measurements. Using the LASBI system, a fiber optics link can be guarantied to be strain free, which is

the key information for its long-term reliability. Tests have been conducted on currently operating

telecom cables and show that local strain can be introduced in the fiber by the installation procedure.

The different fiber sections of fiber optics cables can also be easily identified, which is an interesting

feature for diagnostic purposes. The results from this technique are steadily improving along with

system optimization, in particular for spatial resolution. Other measurements were carried out on

different installations such as aerial cables. These measurements show very minor fluctuations of the

Brillouin shift, so that fibers are granted to be strain free. References

1. T.Horiguchi, T.Kurashima, M.Tateda, IEEE Photonics Technol. Lett., 2, (1990), p.352.

2. M.Tateda, T.Horiguchi, T.Kurashima, K.lshihara, /. Lightwave TechnoL, 8, ( 1990), p. 1269.

3. L. Thevenaz, M. Nikles, P. Robert, in Symposium on Optical Fiber Measurement, Boulder CO,

NIST Special Publication 864 (US Gov. Printing Office, Washington, D.C.), p.211 (1994),

4. M.Nikles, L.Thevenaz, P. Robert, Optics Lett., 21, (1996), p.758.

5. M. Nikles, C. Gabioud, L. Thevenaz, P. Robert, in 3rd Optical Fibre Measurements Conference , OFMC'95 (Liege), paper VI1.2. (1995).

114
Non-invasive mapping of dispersion in optical fibers

R. M. Jopson

Crawford Hill Lab

Bell Labs, Lucent Technologies

Holmdel, NJ 07733 USA

M. Eiselt

Heinrich Hertz Institut

Berlin, Germany

Fiber chromatic dispersion has played an important role in the design of optical

fiber systems for more than a decade. Until the advent of the eminently practical erbiumdoped fiber amplifier, the systems were more or less linear; hence, it was only the

integrated dispersion over a fiber span that influenced system performance. As we have

become more sophisticated in our use of optical fiber, and ever more greedy in extracting

performance from lightwave systems, fiber nonlinearity has become important either as a

tool or as an impairment. As a consequence, it has become important to know the

distribution of the chromatic dispersion along a fiber span rather than just the average

dispersion. Dispersion management, in which a dispersion "map" is chosen so as to minimize the effects of fiber nonlinearities, is becoming an important tool.

Traditional fiber dispersion measurement determines only the average dispersion

of a fiber span. However, in dispersion-shifted fiber (DSF), dispersion is known to vary as a function of location in the fiber. In one (destructive) measurement[l] using a 10-km

length of DSF cut into four 2.5-km segments, it was found that the average dispersion zero wavelength for the segments varied by at least 1 nm - a significant deviation for some

applications. Deployment of erbium-doped fiber amplifiers (EDFA's)in lightwave systems

has led to "dispersion management" as a tool for system designers to further enhance

system performance and to suppress optical nonlinearities.

An early method[2] for mapping the distribution of dispersion in a fiber span is based on Rayleigh backscattering. This technique relies on the dependence of the

dispersion zero on the fiber core size. If the doping of the preform does not change over its length then changes in the dispersion zero can be inferred from changes in the core size.

Changes in core size can be estimated using optical time-domain reflectometry (OTDR) to determine the capture ration for Rayleigh backscattered light. By summing OTDR

measurements taken in opposite directions, the effects of fiber loss can be removed and

changes in the capture ratio can be observed and used to determine variation in the fiber

dispersion zero.

Recently, a non-destructive dispersion measurement method was described[3] that

determined the local dispersion zero from modulation-instability-induced gain at wavelengths longer than the dispersion zero. A strong pump pulse of wavelength Xp and a weak signal pulse of wavelength A,^ are injected simultaneously into a test fiber with Xp-X^

being 5 to 10 nm. The backscattered signal light is observed through OTDR. When the

115
pump wavelength is near the dispersion zero, but in the anomalous dispersion region, the

modulation instability will provide gain for the probe pulse, gain that can be observed in

the OTDR trace. Thus reduction in the slope of the OTDR at a particular distance into the

fiber indicates that the pump is experiencing anomalous dispersion at that point in the

fiber. To map fiber dispersion, the pump and probe wavelengths are swept, maintaining a constant separation, Xp-X^, and the resultant OTDR traces are recorded. The dispersion zero of a particular point in the fiber is at the short-wavelength side of those pump

wavelengths for which modulation-instability gain is observed. This technique has

demonstrated a wavelength resolution of 0.2 nm and a distance resolution of about a

kilometer.

Another technique[4] uses partially-degenerate four-photon mixing[5,6] to determine the dispersion zero. This technique will be the topic of the remainder of this

summary. The mixing generates an idler wave from pump and signal waves of angular

frequencies co^ and (O^ propagating in the fiber. The power of the idler wave with

frequency CO, = 2cOp -co^, will be maximized when the process is phase matched, that is, when Ap = 2P(cOp)-p(0L)J-p(cO/) = 0, where p(co) is the propagation constant. To first approximation, phase matching occurs when co^ is set to the dispersion zero of the fiber[5].

We measured the zero-dispersion wavelength by tuning co^ and looking for a maximum in

the idler power. Distance resolution was obtained by using signal and pump pulses with

widely-separated wavelengths. As illustrated in Figure 1, the differing group velocities of

the pump and signal pulses cause the pump pulse to overtake the signal pulse (assuming

the pump wavelength is near the dispersion zero of the fiber). With sufficient groupvelocity dispersion and short enough pulses, the region of overlap of the pulses occurs over some useful distance which will be the distance resolution of the technique. The timing of

the pulses at co^ and (0^ can then be adjusted so that this overlap occurs at some desired

point within the fiber. The apparatus used is shown in Figure 2. A master oscillator - power amplifier

chip operating at a single frequency of 1486.45 nm provided the signal. It was modulated

by a LiNb03 Mach-Zehnder modulator to generate a 100-MHz train of pulses having a FWHM of 100 ps. Pump pulses with a FWHM of 60 ps were generated by an actively

modelocked EDFA ring laser which could be tuned over the gain bandwidth of the EDFA

with a combination of a 4-nm FWHM thin-film tunable filter and a 0.1 -nm FWHM tunable

fiber Fabry-Perot filter. The two pulses were combined with a 1480/1550 wavelengthdivision multiplexer and coupled into the test fiber. Relative pulse delay was adjusted by a

variable electrical delay line. The dispersion between the signal and the pump was 140

ps/km. With appropriate delay, the pulses overlapped for about 1 km in the fiber. At the

input to the test fiber, the peak signal power was 0 dBm while the pump had a peak power

of +17 dBm. The generated idler near 1610 nm was observed on an optical spectrum

analyzer. A travelling-wave Bragg grating band-reject filter attenuated the ~1546-nm

signal entering the optical spectrum analyzer to reduce scattering-induced background

within the instrument. The amount of ~1610-nm light generated varied by more than 10 dB

with changes in pump and signal polarizations. This variation was reduced to less than 0.2

dB by modulating the polarization of the signal with a great-circle polarization scrambler.

The spectral features obtained using the scrambler closely mimicked those obtained by

116
optimizing the polarizations at each wavelength.

While the technique requires that co^ be scanned near the dispersion zero of the

fiber, almost any signal wavelength will work. The pulse overlap range in the fiber

decreases as the minus two power of co^ - co^, so a large separation will provide good

position resolution. However, improved resolution comes at the expense of decreased

mixing efficiency. Also, separations beyond the 60-nm used in our measurements put the

mixing-product in a spectral region of decreasing detector sensitivity and increasing fiber

loss. We chose 1485-nm as a convenient compromise for the signal wavelength because

of the availability of high-power sources. Positional resolution can also be improved by

using short pulses, but wavelength resolution will suffer for signal pulse widths much

shorter than 30 ps due to spectral broadening.

Maximal mixing efficiency actually occurs when co^ is set slightly off the

dispersion zero because of Kerr-effect shifts[5] and high-order curvature[71 in P(co). For

our power levels, the Kerr-effect shift was negligible. To determine the higher-order

curvature shift precisely, one needs to know the fourth (and higher) order curvature of

P(co) at each desired location in the fiber. However, we found that the average curvature

obtained from a conventional chromatic dispersion measurement was sufficient to determine the curvature shift to within ±0.1 nm at any location in the fiber. In the

measurement described here, the dispersion zero was the phase-matching wavelength plus

1.2 nm. To test the accuracy of the technique, we chose a 9.7 km length of DSF with an

interesting dispersion-zero profile and nondestructively measured the dispersion zero from

each end. We then cut the fiber into 500-m sections and measured the average dispersion

of each section using conventional means. The results are shown in Figure 3, where the

non-destructive data has been corrected for the P(co)-curvature shift described above by

adding 1 .2 nm to the measured phase-matching wavelengths. This correction was obtained

from a dispersion measurement of the intact fiber. The good agreement between

nondestructive measurements made from opposite ends shows that there are no large

systematic errors that depend on the position of the overlap region. It can also be seen that

the non-destructive measurements agree with the destructive measurements to within ±0.2

nm when the latter data is averaged over 1 km. Some of the observed discrepancy is caused by variation of the dispersion zero within the overlap region.

The spectrum produced by the technique can reveal more detailed information

about variations in the dispersion zero than would be inferred from the 1-km location

resolution. For example, a splice between fibers with different dispersion zeros appears as two peaks in the mixing spectrum. One of these peaks grows as the delay increases while

the other peak shrinks. By observing the growth and diminution of these peaks, one can

obtain a more accurate fiber dispersion-zero map than is obtained when this information is reduced to the one number shown in Figure 3. Figure 4 shows the noninvasively-measured

map of the zero-dispersion wavelength of a concatenated fiber. Analysis of the spectral

peaks in the region surrounding the fiber splice near 24 km, shows that the distance

resolution (15% to 85% of asymptotic peak height) of the measurement shown here is 1.4

km.

117
The growing awareness of a need for non-destructively mapping chromatic

dispersion in fibers has spawned a number of techniques for measuring zero-dispersion

wavelengths or in some cases, the measuring fiber dispersion at some desired wavelength.

Current non-invasive techniques for dispersion-zero measurements have a distance

resolution of 0.1 to 1 km and a wavelength resolution of several tenths of nanometer.

Advances are being made rapidly, and it is likely that before this talk is given, we will see new techniques and resolution improvements to existing techniques.

[1] K.Inoue, "Four-wave mixing in an optical fiber in the zero-dispersion wavelength

region", J. Lightwave TechnoL, 10, pp. 1553-1561, (1992).

[2] M. Ohashi and M. Tateda, "Novel Technique for measuring longitudinal chromatic

dispersion distribution in singlemode fibres". Electron. Lett., 29, 426-428, (1993).

[3] S. Nishi and M. Saruwatari, "Technique for measuring the distributed zero-dispersion

wavelength of optical fibres using pulse amplification caused by modulation

instability". Electron. Lett., 31, pp. 225-226, (1995).

[4] A. M. Vengsarkar, and U. Koren, "Non-destructive dispersion -zero measurements

along an optical fibre". Electron. Lett., 31, pp. 21 15-2117, (1995).

[5] R. H. Stolen and J. E. Bjorkholm, "Parametric amplification and frequency

conversion in optical fibers", IEEE J. Quantum. Electron., QE-18, pp. 1062-1072,

(1982).

[6] N. Shibata, R. P. Braun, and R. G. Waarts, "Phase-mismatch dependence of efficiency

of wave generation through four-wave mixing in a single-mode optical fiber," IEEE J. Quantum. Electron., QE-23, pp. 1205-1210, (1987); see also R. W. Tkach, A. R.

Chraplyvy, F. Forghieri, A. H. Gnauck, and R. M. Derosier, "Four-Photon Mixing

and High-Speed WDM Systems", J. Lightwave TechnoL, 13, pp. 841-849, (1995).

[7] C. Lin, W. A. Reed, A. D. Pearson and H-T. Shang, "Phase matching in the

minimum-chromatic-dispersion region of single-mode fibers for stimulated fourphoton mixing," Opt. Lett., 6, pp. 493-495, (1981).

118
Figure 1: Schematic of the

principle used in the four- wave mixing method for mapping fiber dispersion

zeros. The relative positions

of the pump pulse, labeled

"p", and the signal pulse,

labeled "s" are shown for successive times, t, as the

pump overtakes the signal.

During the overlap, an idler pulse is generated which will

travel at a group velocity

close to that of the signal

pulse.

pump

Position in fiber

EOFA BPF FP PC

cm

MZ

MOD

100-ps pulses

9 100 MHz

MOPA

4 nm 0.1 nm

PC POL eops (Yin i-TiAA

variable delay

TEST

"\ SPOOL

MZMOD

BRF

Figure 2: Apparatus. BPF, tunable band-pass filter; BRF. band-reject filter; EDFA, erbium-doped

fiber amplifier; FP, tunable Fabry-Perot filter; MOPA, master oscillator power amplifier; MZMOD, Mach-Zehnder modulator; OSA, optical spectrum analyser; PC, polarization controller; POL, linear polarizer; POL SCRAMBLER, polarization scrambler; WDM, wavelength-division

coupler; XCAV, external-cavity laser.

119
J

I I I I I I I 1 I L 0 5 10 Position (km)

Figure 3: Measured dispersion zero as a function of position in a 9.7-km spool of fiber. The

circles and triangles show the results of noninvasive measurement from either end of the fiber. The

position resolution was about 1 km. The squares show the zeros in the dispersion curves measured

by cutting the fiber into 500-m segments. The error bar shows the ±0.1 -nm uncertainty of the

destructive technique.

E 1554

c

N

0 10 20 30

Position in fiber, km

Figure 4: Dispersion-zero wavelength map of a 33.6-km fiber consisting of 3 concatenated pieces. Horizontal lines show the "average" dispersion zero of each piece as measured by

conventional means.

120
Determination of Zero-Dispersion Wavelength

in Optical Fiber Using Four-Wave Mixing

J. B. Schlager, S. E. Mechels, and D. L. Franzen

National Institute of Standards and Technology, 325 Broadway, Boulder CO 80303

Abstract

Wavelengths of measured maximum four-wave mixing efficiency in optical fiber are compared to zero-dispersion wavelengths measured with a highly accurate frequency-domain

phase shift technique. The average absolute discrepancy between the two wavelengths

determined on fifteen fibers is 0.19 nm; for an average pump-probe spacing of 5.9 nm, the average

spectral width of the four-wave mixing efficiency curve is 0.45 nm.

I. Introduction

Four-wave mixing (FWM) is a parametric process that involves the interaction of four

optical waves in a material.' The efficiency of this process depends on the nonlinear

susceptibilities of the material, the light intensities, and phase matching between the light waves.

In optical fiber optimal phase matching for FWM occurs at or near the zero-dispersion

wavelength.^ In this letter, we describe measurements of partially degenerate four-wave mixing

efficiency near the zero-dispersion wavelength of several different fibers. In partially degenerate

four-wave mixing the pump provides two of the four optical waves at the same optical frequency.

The other two waves are the probe and the generated anti-Stokes light. With the probe light fixed

at a given wavelength, we determine the pump wavelength which gives maximum anti-Stokes

light intensity in the zero-dispersion wavelength region of the fiber. We compare this pump

wavelength with accurate measurements of the zero-dispersion wavelength using the

fi"equency-domain phase-shift technique.^

'*

II. Experiment

The experimental configuration for obtaining FWM efficiency is shown in Fig. 1. Two

tunable lasers provide the pump and probe lights. The pump laser is amplified with an erbium-doped fiber amplifier (EDFA) to a power of +10 dBm. The probe laser light at -3 dBm is mechanically chopped to allow narrow-band detection of the FWM signal with lock-in amplifier

techniques. Light from each source passes through an optical bandpass filter to minimize the

amplified spontaneous emission (ASE) light, which can overwhelm the FWM signal, and

polarization controllers are used to bring the polarization states into coincidence. The pump and

probe lights are combined in a fijsed-fiber coupler, traverse the test fiber, and continue on to the

photoreceiver. An additional optical bandpass filter with sufficient bandwidth (2 nm, full width at half maximum) for the wavelength scan is used at the photoreceiver to fijrther isolate the FWM

signal. The detected signal is passed to a lock-in amplifier. A computer automatically tunes the

pump laser and acquires the FWM efficiency data. Laser wavelengths are measured with an

interferometric wavemeter having an accuracy of 1 ppm.

121
Figure 2 shows the frequency-domain phase shift (FDPS) system, which has been

described elsewhere.^-"* The output of a tunable laser is intensity modulated by an integrated-optic

Mach-Zehnder modulator. The phase of the RF modulation is monitored by a vector voltmeter

after detection, amplification and filtering. The test fiber is temperature controlled to ±0.15 °C.

Laser wavelength is monitored by a wavemeter and matched to corresponding relative group

delays with a temporal resolution of 0. 15 ps. This system can determine zero-dispersion

wavelength with a precision of better than 0.025 nm (three standard deviations).^

in. Results

A typical FWM spectrum is shown in Fig. 3. We acquired this spectrum with an optical

spectrum analyzer placed just after a 10 km dispersion-shifted test fiber while the pump was tuned

to the fiber's zero-dispersion wavelength. During FWM efficiency measurements, the probe light

exiting the test fiber is attenuated relative to the FWM light by the additional optical bandpass

filter. This ensures good FWM signal isolation at the lock-in amplifier. A FWM efficiency curve of a 10 km dispersion-shifted fiber is shown in Fig. 4. Here the mid-scan pump-probe wavelength

separation was about 5 nm and the efficiency curve had a fijll width at half maximum (FWHM)

width of 0.36 nm. The experimental data are plotted with a theoretical curve.^ Although the

match is good, we believe the discrepancy between the theory and experiment is primarily due to the variation of zero-dispersion wavelength along the length of the fiber. Some test fibers

exhibited more complicated FWM efficiency curves, but a clear FWM efficiency peak was still discernible in these fibers. In addition, by increasing the pump-probe wavelength separation, we

were able to achieve narrower efficiency curves. A mid-scan pump-probe wavelength separation

of 1 1 nm (the maximum possible with the filters available) gave a FWM efficiency curve width of

0.11 nm FWHM with no appreciable shift in peak FWM efficiency wavelength. Theoretically, the

maximum efficiency wavelength does not depend on pump-probe separation; however, in practice

there are trade-offs. A small separation results in a spectrally broad efficiency curve while a large

separation enhances the unwanted effects of polarization mode dispersion. The 6 nm mid-scan

separation chosen for most of these measurements represents a good compromise given the

optical bandpass filters available to us.

Comparisons between peak FWM efficiency wavelengths and zero-dispersion wavelengths

measured by the FDPS system for fifteen commercially available dispersion-shifted fibers are presented in Table I.

In all fifteen fibers a single FWM efficiency peak clearly dominated, and the absolute

discrepancy between this peak and Xo was on average less than 0. 19 nm. The mid-scan

pump-probe wavelength separation was about 4-5 nm for fibers A-D, 1 1 nm for fiber E, and 6 nm

for fibers F-0. The dispersion slope dD/dX measured with the FDPS system ranged from 0.069

to 0.0735 ps/(nm^.km) with an average of 0.0713 ps/(nm^. km). The measured PMD at 1550 nm

ranged from 0.03 to 0. 1 1 ps/.ykm for fibers A-E, and varied as much as 30 % for each fiber

measured over time.

122
Most fiber specimens were examples of recently manufactured high-quality fibers. Cut-back measurements on four equal-length sections of one fiber showed a linear increase in zero-dispersion wavelength of 1.5 nm over a length of 10 km.

TABLE I Comparison of Peak FWM Efficiency and Zero-Dispersion Wavelength

Fiber (length) X be ^ FWM Spec. Width (nm) (nm) (nm) (nm, FWHM) (nm)

A (5 km) 1550.27 1544.69 1544.68 0.44 0.01

B (12 km) 1555.21 1551.36 1551.38 0.48 0.02

C (12 km) 1556.40 1552.54 1552.70 0.41 0.16

D(10 km) 1558.26 1553.37 1553.45 0.39 0.08

E(lOkm) 1560.14 1549.20 1549.21 0.11 0.01

F (20 km) 1555.14 1548.94 1549.17 0.21 0.23

G(lOkm) 1555.14 1548.96 1548.86 0.32 0.10

H(lOkm) 1555.14 1549.14 1549.44 0.36 0.30

1(5 km) 1555.14 1549.12 1549.04 0.51 0.08

J (5 km) 1555.14 1549.57 1549.87 0.35 0.30

K (2.5 km) 1555.14 1549.12 1548.83 0.78 0.29

L (2.5 km) 1555.14. 1549.14 1549.22 0.67 0.08

M (2.5 km) 1555.14 1549.33 1549.56 0.55 0.23

N(2.5 km) 1555.14 1549.44 1549.94 0.48 0.50

0(1.25 km) 1555.14 1549.37 1549.82 0.67 0.45

IV. References

1. G. P. Agrawal, Nonlinear Fiber Optics, 2nd Edition. New York: Academic Press, 1995.

2. K. Inoue, "Four-wave mixing in an optical fiber in the zero-dispersion wavelength region,"

J. Lightwave Technol, vol. 10, no. 11, pp. 1553-1561, 1992.

3. S. E. Mechels and D. L. Franzen, "Accurate measurements of zero-dispersion wavelength

in single-mode fibers," 3rd Optical Fiber Measurement Conference Technical

Program, paper IV.2, 1995.

4. S. E. Mechels, J. B. Schlager, and D. L. Franzen, "Accurate zero-dispersion wavelength

measurements in single-mode fibers: two frequency-domain methods," LEOS 1995

Meeting Proceedings, vol. 2, pp. 75-76, 1995.

5. S. E. Mechels, J. B. Schlager, and D. L. Franzen, "Accurate measurements of

zero-dispersion wavelength in optical fibers," to be published in J. Res. Nat 7. Inst. Stands. Technol.

Contribution of NIST, not subject to copyright.

123
Tunable laser (Pump)

PC

nnn

Optical bandpass

niter

EDFA

50/50

Fuscd-flber coupler

Tunable laser (Probe)

Choppei

PC

OOP

Optical bandpass

niter

Chopper

controller

CPU

Reference

Lock-in ampUner

signal

Photoreceiver

Optical bandpass

niter

Figure 1 . Experimental configuration for obtaining FWM efficiency:

PC, polarization controller; CPU, personal computer.

Mach-Zehnder

Tunable Laser Modulator

Crystal Oscillator

at 2 GHz

Wavemeter

CPU

Fiber Sample

Vector Voltmeter

Photo- Narrow Band Filter Receiver

at 2 GHz

Figure 2. Frequency-domain phase shift system: CPU, personal computer.

Punp

Wavelength (nm) Figure 3. Partially degenerate FWM spectrum:

FWM, four-wave mixing anti-Stokes peak.

Pump Wavetengffi (nm) Figure 4. FWM efficiency versus wavelength

in 10 km dispersion-shifted fiber. Mid-scan

pump-probe separation is 7 nm.

124
Accuracy issues in comparisons of time- and frequency-domain

polarization mode dispersion measurements

Optoelectronics Division - National Institute ofStandards and Technology

325 Broadway, Boulder CO 80303, U.S.A.

If experimental comparisons are to be made between time- and frequency-domain measurements of

polarization mode dispersion PMD, they must be done with a good understanding of the systematic

and random uncertainties present. In this paper, I quantify sources of systematic uncertainties in time

domain PMD measurements and demonstrate correction techniques on experimental data.

Until recently, theoretical and experimental work [1-10] claimed to support an equality between

PMD measured in the time domain as the square root of the second moment a of the autocorrelation

function and PMD measured in the frequency domain as the RMS differential group delay (DGD),

so that RMS DGD/o = 1. However, the recent theory of Heffiier [1 1] shows the issue to be more

complicated, with RMS DGD/a depending on the spectral shape of the time domain measurement

source and only for large values of PMD-bandwidth product does the ratio demonstrate a constant

value 0.866 which disagrees with previous theory by about 13%. Most of the referenced attempts

to measure the RMS DGD/o ratio lacked the precision to see a 13% effect. Furthermore, standard

second moment evaluations of time-domain PMD have several sources of systematic uncertainties

large enough to nullify comparison efforts. These systematic errors were avoided by one recent

experimental work [12] which used a curve-fitting method to measure o giving a value of RMS

DGD/o close to 0.9. However, curve-fitting is not always considered to be rigorous enough. So, I present corrections here which allow significant reductions of the systematic biases of the standard

second moment calculation.

Measurements of PMD in highly polarization-mode-coupled optical fibers using an interferometric or equivalently wavelength scanning with Fourier transform (WSFT) technique produce an

autocorrelation function illustrated by the one-sided example of Figure 1. This time domain response

is a quasi-random amplitude (due to random phasing) under a Gaussian envelope. The PMD of the

fiber can be measured by finding the square

P.A. Williams

However, in calculating this quantity,

systematic errors due to numerical

integration, data clipping and central peak

removal are encountered and must be

corrected.

0 4 8 12 Time Delay (ps)

16 20

Figure 1 Typical Fourier-transformed

wavelength scanning data.

!

125
A good approximation to Equation (1) is obtained using the trapezoidal integral approximation

a

1(1. )tl --i(t

(2)

k = a.

where t^ and tb are the time coordinates of the first and last points used. A simpler rectangular

approximation is sometimes used [13,14]. However, the numerical integration is usually carried out

with the time coordinate sampled at the practical values t = 0, A, 2A,... for example, as opposed to the more appropriate t = A/2, 3A/2, 5A/2,... values (A is the sampling step size). It should be

cautioned that for a single-sided data set as from WSFT, integrating from ta=0 will underestimate

o by an amount dependent on the time domain sampling density r\ (number of data samples / a).

Under practical conditions, this can result in errors of several percent and the trapezoidal

approximation represents a significant improvement (Figure 2).

Another systematic bias comes from limited

range. The sampling range tt, can be limited by

the wavelength resolution in WSFT

measurements or the scan range of the

interferometer. But, the first limit to be

encountered is likely to be due to the signal to noise ratio of the system. Generally, the useable

portion of the data is determined by some

criteria such as the exclusion of data beyond the

point tb where the signal drops below 2 times the rms noise level [13,14]. However, clipping the

data wing(s) in this way systematically biases

the measured a toward lower values. This can be seen from Equation (1) in that for a nominal

gaussian, at large values of t, the positive

integrand of the denominator I(t) approaches

zero more quickly than the positive integrand of

the numerator I(t)t^. So, the result of excluding

data at large values of t will be to bias o toward

smaller values. Figure 3 shows a computer

simulated result of measurement error as a function of the data clipping point tjo (solid

curve) with a being the true value of the square root of the second moment. The clipping error is substantial. Indeed, an RMS noise level of 6 or 7% would require clipping at t^/o = 2 giving an error on the order of 10% - large enough to bias measurements of RMS DGD/a from 0.866 up to 1.0.

1.00

0.98

0.96

0.94

0.92

0.90

-1

1 I I I I I 11 1

1 I I I I I M 11 I I I I I U

I I I I I I I 1

I I I I I I I 1

r I I I I r I1000

Measurement errors for trapezoidal

and rectangular (solid) integral

approximations as a function of sampling density

(the number of sampled points/a) r|.

Figure 2

(dashed)

126
1.0 To correct for this error, a simple look-up table

is useful. The dashed curve of Figure 3 is the

solid curve plotted as a function of t^/o^ where

is the measured value of second moment. In

practice, dividing by the value of the dashed

correction curve at t^/o^ removes the clipping

error. Of course, the correction has trouble for

values of t^Jo^ which are much less than 2 where

the curve becomes very steep and small random

errors in measuring are amplified by the

correction factor. An example of the correction

performance is shown in Figure 4 where a is measured on simulated WSFT data (using the

simulation of Reference 1 1) at various clipping

levels and then compensated to give a corrected

value. Clearly, the compensated a performs well,

within 2% of its undipped value down to about

the tjo^ = 2 level.

0.0

1 I I I

I

I—I—I I

I

I—I IL I I I—

I—I—I—I—

r

Calculated error - Correction Curve _

I I I I I I I I I I I I I I I I I I I I r I f

Figure 3 Data clipping error vs. sampled fraction

oftrue second moment (solid) or measured second

moment (dashed).

1.1

1.0-

,

0.9 -

0.8

0.7

0.6

0.5

T—I—I—I—I—I—I—I—I—

I—I—I—1—I—I—I—I—I—

r

Corrected Uncorrected

- i -

!l

Another source of systematic error comes from

removing the central peak from the time domain

response. In autocorrelation interferometers, the

interferogram has a large source autocorrelation

peak around t =0 which contains no PMD

information. A similar peak results from

windowing in WSFT. In practice, o is often

calculated by simply adjusting tg in Equation 2 to skip the data points under the central peak.

However, the result is a systematic increase in

measured o. The reason can be seen in Equation

2, where eliminating low t values of I(t) reduces

the value of the denominator more drastically

than does eliminating low t values of I(t)t^ from

the numerator. The simplest technique to compensate this error is to substitute the values

of I(t) where the central peak occurs with

estimates of what I(t) would be without the

central peak. As an example, this is illustrated for

single-sided WSFT data (the procedure is identical for interferometric data). First, o^, is measured using Equation (2) with set to ignore

values of I(t) which are dominated by the central

peak.. Then is used to define new integration limits t^ = oJ2 and tb = a^. Then 00.5. i is measured

using Equation (2) with these new limits and divided by 0.75 (the value of ao.5., for an ideal Gaussian

of unit amplitude). The result of this division is an estimate at the amplitude A of the Gaussian

J 1 I L

tb/a

Figure 4 Normalized simulated measurements of

uncorrected (solid) and corrected (dashed) second

moment vs. sampled fraction of second moment.

127
function describing the time-domain response

m=Aexp( ). (3)

Then, and A are used along with Equation

(3) to generate estimated values of I(t < t^ ). Finally, o is recalculated using equation (2)

with t=0. An example of the error and

correction results is shown in figure 5 for

computer simulated interferometric data.

It is worth mentioning that there is one

systematic error in time domain

measurements which is well known but often

inappropriately corrected for. I refer to the source width subtraction technique

'True ^Meas ^Source ' (4)

The idea is to subtract the square root of

second moment of the Fourier transform of

the optical source spectrum agource f^om the

measured square root ofthe second moment of

the interferogram o^^zs to yield the true square

root of the second moment of the spectral

response of the device under test Oj^^. This

technique comes from the fact that a ftinction

which is the product of two other functions

has a variance of its Fourier transform which

is equal to the sum of the variances of the

Fourier transforms of the other two functions.

However, application of Equation (4) to a WSFT or interferometric PMD measurement

is inappropriate since OMeas is measured after

the time domain response has undergone a

non-linear enveloping technique. Specifically,

this involves using the envelope of the

photocurrent in the interferometric case or using the magnitude of the Fourier transform

in the WSFT case. Both of these enveloping

functions are non-linear modifications to the

process which make Equation (4) nonT3

O

z

2.2

2.0

1.8

1.6

1.4

1.2

1.0

0.8

Uncorrected

Corrected

0.0 0.2 0.4 0.6

t.

0.8 /a

1.0 1.2 1.4

Figure 5 Normalized simulated measurements of

uncorrected (solid) and corrected (dashed) second

moment vs. starting measurement time/measured

second moment.

1.1

I.O

0.9

0.8

0.7

0.6

-1—I—I—

O

Second Moment (full data set) ^ Second Moment (averaged) A Corrected Second Nfoment (averaged) O Curve Fit (averaged) rill

RMSDGD(ps)

12 16

Figure 6 Ratio of frequency- and time-domain

PMD measurements as a function of PMD.

applicable. So, while a finite source width still contributes a systematic bias to the measured second

moment, there is not yet a rigorous method of correcting the error.

128
In order to demonstrate the usefulness of the above corrections, the value of RMS DGD/a is calculated using the data set from reference [12]. The data represents measurements of seven

different fiber specimens using both Jones matrix eigenanalysis and WSFT techniques, a is extracted

from the WSFT data using Equation 2 with no correction factors. Then, the same o is corrected using

the above techniques. Finally, a third set of second moment values are obtained by performing a nonlinear least squares fit of Equation 3 to the data sets. This curve-fit technique is not as susceptible

to the systematic errors described above and acts as an estimate of the true value of a. In Figure 6,

the ratio of the frequency domain measurement of PMD (RMS DGD) and the associated time

domain measurement a are plotted as a function of the PMD of the fiber (RMS DGD). The upper

dashed line indicates the previously predicted 1.0 ratio and the lower dashed line illustrates the

asymptotic 0.866 value predicted by Heffrier. The large data points represent the average of several

statistically independent measurements. The small diamonds are the unaveraged uncorrected second

moment measurements to illustrate the spread in data and the number of measurements. Clearly, the

uncorrected second moment data yields values with a systematic bias of -5% above the curve fit results. Application of the correction techniques to the data brings the second moment values into

good agreement with the curve fit results as expected. Clearly, the corrected results indicate a ratio

which is closer to 0.866 than to 1.0. However, a larger population of measurements is needed to make a quantitative judgement.

My thanks go to B.Heffher for providing some of the simulated data used in this work andfor

sharing with me some insight into the world ofPMD.

[I] N.Gisin, Opt. Commun., 86, 371-373 (1991).

[2] N.Gisin, and J.P.Pellaux, Optics Communications, 89, 3 16-323 (1992).

[3] B.Pemy, C.Zimmer, F.Prieto, and N.Gisin, Electron. Lett., 32, 680-681 (1996).

[4] M.Artiglia, A.Chiantore, and A.Rossaro, Proc. Optical Fiber Measurements Conference '95, p. 1.4 (1995).

[5] M.Artiglia, P.Morra, and C.Sartori, Proc. Optical Fiber Measurements Conference '95, p. II.6 (1995).

[6] N.Gisin, Proc. Symp. Optical Fiber Meas. '94, 149-154 (1994).

[7] N.Gisin, R.Passy, J.C.Bishoff, and B.Pemy., IEEE Phot. Tech. Lett., 5, 819-821 (1993).

[8] Y.NamihiraandJ.Maeda,Elec. Lett., 28,2265-2266(1992).

[9] Y.Namihira and J. Maeda, Proc. Symp. Optical Fiber Meas. '92, 145-150 (1992)

[10] N.Gisin, R.Passy, B.Pemy, A.Galtarossa, C.Someda, F.Bergamin, M.Schiano, and F.Matera, Electron. Lett., 27,

2292-2293 (1991).

[II] B.L.Heffher, Opt. Lett., 113-115 (1996)

[12] P.A.Williams and P.R.Hemday, Proc. Opt. Fiber Meas. Conference '95, p. 1.2 (1995).

[13] Fiber Optic Test Procedure (FOTP) 113, Telecommunications Industry Association.

[14] Fiber Optic Test Procedure (FOTP) 124, Telecommunications Industry Association.

129

How accurately can one measure a statistical quantity like

Polarization Mode Dispersion?

N. Gisin^ B. Gisin\ J. P. Von der Weid^ and R. Passy^

Group of Applied Physics University of Geneva, 1211 Geneva 4, Switzerland

Center for Telecommunication Studies

Pontifica Universita Catolica de Rio de Janeiro

Rio de Janeiro, Brasil

Different measurement techniques of PMD are presently available and under further investigations [1, 2, 3, 4]. All these techniques involve an averaging procedure. Indeed, this is necessary,

since the Differential Group Delay (DGD) is a statistical quantity and PMD is defined either

as the mean DGD or its RMS value. However, the statistical character of the DGD and the

fact that any measurement involves only a finite sample set imply that there is an intrinsic

limitation to the accuracy of PMD measurements. This limitation is independent of any measurement setup implementation, but is unavoidable, similarly to the well-known example of the

measurement of the probability of coin tossing with only a finite number of trials. The main

difficulty in evaluating this limitation in the case of PMD is that, contrary to the coin tossing

example, the DGD at nearby wavelengths are not independent.

In order to present our result, let us introduce some notations. Let f2(cu) denote the

principal polarization state vector [6]. By convention, its norm equals the DGD: 8t{uj) =

Y^i(^) + ^y{'^) + ^zi^)- practice one is interested in statistical ensembles or in time ensembles. However, wavelength ensembles are much easier to produce. Furthermore, it is generally

admitted that these different ensembles are statistically equivalent and there are good reasons

[7] for doing so. Consequently, all major PMD measurement techniques are based on a wavelength ensemble [9, 10]. Note that large wavelength ranges correspond to long time fluctuations.

Accordingly, we define the PMD Delay Ar as the RMS value of the 6t{uj):

In practice the optical frequency range is limited by the laser tunability. Another limitation

is due to the slight wavelength dependence of the modal birefringence and the polarization

mode coupling length. However, in order to define a PMD value vahd for arbitrary long time

fluctuations, we shall assume in our model that the local modal birefringence and the mean

polarization mode coupling length are wavelength independent. This amounts to assume that

the modal field does not change over the considered wavelength range (i.e. all wavelengt.hs see the same refractive index structure), hence the wavelength dependence of 6r(uj) is only due to

interferences between the different optical trajectories in the fiber [4]. With such an assumption,

the theoretical mean PMD delay At can be defined as:

Defining PMD this way has the great advantage that the probability space is well defined and the

statistical ensemble of interest consists in all optical frequency intervals of length Aa;. Henceforth

(1)

At = lim At(0, u>2) (2)

131
we study the variance of Ar(a;i,u;2)^ when Ao; = uj2 — u)] is finite (recall that the variance is defined as: Var{X) = mean{X'^) — Tnean{X)^)

Let ujj, j=l..N denote the optical frequencies at which the DGD 6T(u)j) are measured. Hence,

this finite sample set corresponds to a measured value of PMD STmeas-

^''Las = ^ E ^^(^j)' = ^I1 {^M? + ^vM' + ^^'Kf) (3)

where each component of the principal state vector f2 is a normal random variable [8] of zero mean and variance At /3. Hence, Ar^p^^ is the sum of three normal random variables of mean

At /3. Recall that for any normal random variable X, Var{X-')='2Var{X)~, and that the

variance of the sum of independent random variables is simply the sum of the variances. Hence,

4 if the 9.k{ujj) would be independent, the variance of Ar^^^^ would be equal to but, clearly,

they are not independent (think for instance that is a continuous function of u, values of

6T{iL)) at interpolating frequencies uj provide thus little additional information). The number of

data that are eflTectively independent can however be estimated as follows. The functions Qf-i^)^

k=x,y,z, have irregular oscillations, hence using the Nyquist theorem two points per cycle can be considered as independent. But the number of extrema of ^ki^) related to the PMD

Ar, similarly to the well known case of the extrema of the intensity passing through a fixed

analyzer [3]. Hence, we arrive at our main result:

/ 9 \ IStt 2/c7r At At' , .

Consequently, a measurement of the DGD over a wavelength range Acj provides an approximate

value ATmeas of the theoretical mean PMD delay At:

ATrnen,^A^(l±--^^] (5) V VAtAljJ

where a = ^ 0.76. A more elaborated derivation [12] of relation (5) and our numerical

simulation confirm this relation, but with a ^ 0.9. For example, if Au) = 100— (corresponding in wavelengths to 6 A 130 nm around 1550 nm

or to 6\ ^ 90 nm around 1300 nm), then for a PMD of 1 ps, the measurement uncertainty due

to the statistical nature of the DGD is of about 9%! For a PMD of 0.1 ps, the uncertainty is even as large as 28%.

It .should be stre.ssed that the above limitation (5) is not specific to one measurement scheme,

but is intrinsic to all measurement techniques that average the DGD over a wavelength range.

For example the figures 1 to 3 illustrate our result on simulations. Figure 1 presents on 5000

numerical simulations the case of the Jones Matrix Eigenanalyses (JME) measurement method

simulating a tunable laser over 90 nm around 1300 nm. The theoretical mean PMD value is obtained by computing the RMS DGD over 50.000 ^. Each point represents one run of the

simulation software. The thick lines are the mean values and the mean plus and minus one

standard deviation. The thinner lines are the theoretical curves with a = 0.9. Figures 2

and 3 are analogous, but for the extrema counting and interferometric measurement schemes,

respectively. Figure 4 shows the ratio of simulated measurement results as obtained with the

JME and interferometric methods.

In conclusion, a general formula (5) of the accuracy limitation to PMD measurement due to the statistical nature of the DGD has been derived. Present day commercial instruments are

132
already close to the predicted theoretical accuracy limitation.

Acknowledgment Financial support by the Swiss "Office Federal de I'Education et de la Science"

within the European project COST 241 is gratefully acknowledged. Thanks are also due to Dr B. Perny

for many helpful discussions.

References

[1

[2

[3

[4

[5

[6

[7;

[8,

[9

[10

[11

[12

N. Gisin, J.P. Pellaux and J. P. Von Der Weid, J. Lightwave Tech. 9, 821-827, 1991.

B.L. Heffner, IEEE Photon. Tech. Lett. 4, 1066-1069, 1992.

CD. Poole and D.L. Favin, J. Lightwave Tech. 12, 917-929, 1994.

N. Gisin and the COST 241 Group, Pure and Applied Optics 4, 511-522, 1995.

B. Perny, Ch. Zimmer, F. Prieto and N. Gisin, Electron. Lett. 32, 680-681, 1996. CD. Poole and R.E. Wagner, Electron. Lett. 22, 1029-1030, 1986.

N. Gisin, Opt. Communications 86, .371-.373, 1991.

F. Curti, B. Daino, G. de Marchis and F. Matera, IEEE J. Lightwave Tech. 8, 1162-1166, 1990.

EIA/TIA, FOTP-113, FOTP-122 and FOTP-124, proposals on PMD measurement methods.

ITU-T, SG15-WP4, Contributions on measurement methods of PMD.

B.L. Heffner, Optics Lett. 21, 113-115, 1996.

N. Gisin et al., preprint, University of Geneva, 1996.

1 Figure Captions

1. Illustration of our main result (5) on 5000 numerical simulations for the case of the Jones Matrix

Eigenanalyses (JME) megisurement method simulating a tunable laser over 90 nm around 1300 nm. Each point represents one run of the software, simulating a concatenation of 200 polarization

maintaining fibers of random orientation and birefringence. The theoretical mean PMD value At is obtained by computing the RMD DGD over 50.000 —. Atjme is computed according to published ps procedures [2]. The thick lines are the computed mean values and the mean plus and minus one

standard deviation. The thinner lines are the theoretical curves (5) with a — 0.9.

2. Same as figure 1, but for the extrema counting measurement method (using the polarization mode

coupling k = 0.82 and the same value for a = 0.9).

3. Same as figure 1, but for the interferometric measurement method (using the second moment of

the interferogram with Heffner's [11] factor and the same value for a = 0.9).

4. Comparison of the JME and interferometric measurements over the same wavelength range. Note

that the standard deviation of this ratio is identical to that of the previous graphs (the thin lines are the same on all graphs). This illustrates that the uncertainty is similar for all measurement

methods.

133
134
Compensation formula for noise threshold bias of

interferometric PMD measurement

B. L. Heffner

Hewlett-Packard Laboratories, 3500 Deer Creek Road, Palo Alto, CA 94304

Interferometric measurement of the polarization-mode dispersion (PMD) of a highly

mode-coupled fiber results in a roughly Gaussian envelope modulated by speckle noise.

The width of the Gaussian envelope is often determined by calculation of the second

moment ofthe interferometric response about the zero-delay point. The noise threshold

required by this approach biases the result, leading to an apparent width smaller than

the Gaussian width that would be measured in the absence of noise. The cause of the

noise threshold bias is examined, and a formula is introduced that allows estimation

ofa noise-free Gaussian width based on real measurements with noise. An ensemble of

experimental PMD measurements demonstrates that this formula removes the noise

threshold bias.

The unprocessed output of an interferometric PMD measurement is the envelope

£(0 of the a.c. photocurrent caused by interference fringes, as a function of the differential time delay t through an interferometer. When a highly mode-coupled fiber is

measured, the resulting envelope £(0 consists of a roughly Gaussian envelope multiplied by a random amplitude analogous to optical speckle. Calculation of the PMD

involves assessing the half-width a of the Gaussian envelope £(r) . The half-width is

commonly assessed by calculating the second moment of £(r) about the zero-delay

point, resulting in a second-moment estimate :

The motivation for (1) can be clarified by the example of a perfect Gaussian G(t)

of arbitrary amplitude and half-width a

:

When £(0 = G{t) , the half-width is found correctly (a^^ = a) by application of (1).

However, an interferometric measurement contains both additive noise and a random

amplitude caused by speckle, whereas G(t) does not. To accommodate the noise we

specify a noise threshold and set the limits of integration of (1) equal to T- and T+ , the points that bound the region in which £(f) exceeds the noise threshold (Fig. 1).

Integrating over the finite domain defined by the intersection of £(?) with the

noise threshold, we obtain a biased estimate a of the half-width:

a

2SM (1)

(2)

y.2

a = (3)

135
Interferometric envelope £(?)

Fig. 1. a.c. photocurrent envelope resulting from an interferometric measurement of

PMD. The second moment is calculated by integration over a central region where the

noise does not dominate the signal. The bounds of this region can be defined by the furthest

points of intersection between the envelope and a noise threshold.

We expect T- ~ T+ so for purposes of correction we'll use the mean value

T = (T- + T+) /2 . No attempt will be made to correct for errors induced by the central

coherence peak, which can be suppressed. If we assume both the noise and the

random amplitude are stationary, and assume the underlying envelope is truly Gaussian, then we can calculate a correction factor z = which will allow the systematic bias caused by the noise threshold to be removed.

where x = -a

Once an envelope £(?) is measured, T is determined by the noise threshold and

§ is calculated from (3). Given T and a , in principle (4) can be solved for a or ^ , but

the author is not aware of such a solution in closed form. Instead, consider the

following approximation for z in terms of the measurable quantity y = ^ - 'Js :

a ^ a

0.969 +

2:2653543 _

0.00023788 ^ q 533438 Exp(- 5.30178 y + 1.09896 /)

for T > 3.85 a

for 1.743a< r < 3.85a (5)

where y = ^-j3

The correction factor z and its approximation z are plotted in Fig. 2 as a function of r/a , the number of corrected half-widths above the noise on either side of the

zero delay point. The fractional error of the approximation (5) is bz = (z- z) /z- This

error is calculated from (4) and (5) for various values of T/o (Fig. 3).

136
1.5 2 2.5 3 3.5 4

Fig. 2. Correction factor z and its approximation ? as a function of T/o . Points are exact

corrections, soKd line is the approximation.

Fig. 3. Fractional error of the correction factor given by (5). Note the difference in the in- dependent variable between Figs. 2 and 3. The ratio T/6 approaches a limit of Js ~ 1.732

for very small integration intervals, i.e. as T/a approaches zero.

From Fig. 3, the factional error 5z between (4) and (5) is seen to be given by

|5z| <0.01 for T> 1.870 a

(6)

|6z| <0.026 for T> 1.743$

Finally, we demonstrate the use of this correction factor with an ensemble of

interferometric PMD measurements. A single 6.6-km single mode fiber, loosely

wound on a spool, was measured 25 times using the block diagram of Fig. 3. Statistical independence ofthe 25 measurements was ensured by changing the temperature

of the fiber in 8-K steps or by rearranging the layers of fiber on the spool between

measurements. We assume the rearrangement and temperature changes had no

systematic influence on the PMD of the fiber, and therefore consider the 25 half-width

measurements to comprise a statistical ensemble. The mean of the ensemble distribution was estimated in terms of an 80% confidence interval, i.e. the mean of the

137
distribution lay within the interval with probability 0.8. Confidence intervals were

calculated using noise thresholds 3 dB, 6 dB, 9 dB, and 12 dB above the measured rms

shot noise. In each case, the resulting interval is shown in Fig. 5 with and without

correction by the ratio given by (5).

INTERFEROIVin-ER

1 DATA 1 ENVELOPE

1 RECORDER | DETECTOR PHOTODIODE

Fig. 4. Block diagram of interferometric PMD measurement.

t

Half-width of

envelope (ps)

3.8

3.7

3.6

3.5

3.4

3.3

t

Increasing noise threshold 3dB 6dB 9dB 12 dB

Fig. 5. 80% confidence intervals indicating the mean half-width of the interferometric

Gaussian envelope, calculated using four different noise thresholds 3 dB, 6 dB, 9 dB, and

12 dB above the measured rms shot noise. Calculations for the filled intervals are corrected

by (5); those for the open intervals are not corrected. Correction by (5) removes the influ- ence of the noise threshold over the calculated PMD.

The signal level at the shoulder of the measured Gaussian envelope was approximately 38 dB above the rms shot noise level (a factor of 80 in photocurrent). Use of

different noise thresholds is equivalent to use of a constant noise threshold coupled

with differing amounts of optical loss in the measured fiber. The open intervals of Fig.

5 demonstrate the dependence of the calculated Gaussian half-width on the noise

threshold. The filled intervals, calculated using the correction given by (5) are essentially independent of the noise threshold level.

Summary: Direct evaluation of the second moment of an interferometric PMD

measurement leads to a bias in the calculated half-width a of the Gaussian envelope.

This bias can be removed by the correction factor given by (5). The true Gaussian halfwidth is given by a = az. Efficacy of this correction is shown in Fig. 5, where uncorrected half-widths are seen to depend on the noise threshold, whereas the corrected

half-widths are very nearly equal regardless of the noise threshold.

138
A practical comparison between two different PMD measurement

methods

Bertil Arvidsson, Hans Mickelsson^

Ericsson Cables AB, Hudiksvall, Sweden

Kim Brising

TELIA, Farsta, Sweden

Abstract

A comparison between PMD measurements using the interferometric and the wavelength

scanning technique is presented. The results, arising from measurements on fibres on shipping

spools, fibres on spools with various diameters and on cabled fibres show a reliable

relationship between the two methods. As a conclusion it can be said that a result obtained by

one of the instruments is reproducible on the other.

Introduction

With increasing bit-rates in telecommunication networks PIVID has become a more important

subject in both installed cables and in cables under development. In order to investigate how

different designs and installation techniques impact on the cable PMD a lot of field measurements as well as lab testing has to be carried out. Because we have used different

instruments in lab environment and in field trials, an investigation of the accuracy between the

two instruments were carried out. For a background of definitions and technique see (1). Measurement set-up and choice of fibres and cables

The two instruments were one EG&G instrument, using the wavelength scanning technique

(fixed analyzer) and one GAP instmment, using the interferometric technique. The main

difference between the performance of these two instruments are the possible ranges. The

actual EG&G instrument has a possible PMD-range between 0.05 to 1.6 ps, the wavelength can be choosen in different regions between 1165 and 1688 nm. The GAP-instrument is capable of measuring PMD from approximately 0.1 ps to 30 ps, the wavelength is chosen by

using either the 1300 nm or 1550 nm LED.

For this comparison two 36-fibre ribbon cables and four bare fibres on shipping spools were

chosen. Also five bare fibres wound on spools with different diameters, introducing different

levels of PMD, were included in the comparison (see also 2).

Conditions and theory

All measurements were conducted in a clean room area with a controlled temperature, both

the fibres and the cables were stored in that environment for at least 24 hours before any

measurements took place. To obtain almost identical conditions for the comparative

measurements they were performed in a sequence, first one and then the other instrument.

One problem in interpreting the results is the fact that different wavelength ranges gives

slightly different results. The GAP-instrument uses two LED:s with unknown spectral widths

and the EG&G-instmment uses two LED:s and a monochromator to control the spectral range.

Since PMD has a slight wavelength dependence especially for short fibre lengths, this fact can have some impact on the results. Another matter to consider is the wavelength window where

the measurements shall be performed.

^ Now at Ericsson Telecom AB, Stockholm, Sweden.

139
Why can the PMD-value differ between measurements done by the interferometric and the

wavelength scanning method?

1. The wavelength range in case of the EG&G instrument is well defined due to the

monochromator and the two LED:s used. Uniform spectral weighting is achieved by

normalising the polarimetric data for the source power and fiber loss (3). For the GAP

instrument, the spectral range is determined by the shape of the LED spectrum (normally

Gaussian), so this may lead to non-unifomi spectral weighting.

2. The deployment conditions of the fibres/cables, the degree of mode coupling and intrinsic

birefringence.

3. The statistical nature of PMD gives that the results between different measurements can

vary significantly, up to 30-40% in some extreme cases.

Although the wavelength dependence typically only shows through in the second decimal point

and all specifications seen so far specifies PMD to the first decimal point, the following

methodology was used to amortise the effects of the wavelength dependence. For

measurements with the wavelength scanning technique the best possible wavelength range

was chosen. This means that sometimes in a series of measurements different ranges might

have been used. However, the results are always treated as wavelength independent. For

measurements using the interferometric technique at both wavelengths the results are averaged and taken as the fiber PMD value.

Results

Fibres on shipping spools

The four bare fibres on shipping spools, labelled A, B, C and D were measured 10 times with

the GAP instrument at both 1310 nm and 1550 nm and 7 times with the EG&G instmment at

different wavelength ranges to obtain a knowledge of the repeatability. The fibres were wound

on standard shipping spools, diameter 150 mm, with a winding tension of approximately 30

grams. Table 1 shows the results. They consist of mean values between all wavelengths used

on each instrument. As can be seen no significant differences are present.

Fibre Instrument PMD mean value Length[km]

Jps/Vkm] [ps]

A GAP 0.04 0.20 19

EG&G 0.06 0.25

B GAP 0.06 0.15 6

EG&G 0.05 0.14

C GAP 0.07 0.15 4

EG&G 0.07 0.11

D GAP 0.06 0.31 25

EG&G 0.07 0.28

Table 1. Repeatability measurements on fibres wound on shipping spools.

Fibres on spools with different bend radii

Five spools with radii ranging from 55 mm tc 200 mm were chosen. These spools have been

measured periodically for approximately 6 months with the wavelength scanning technique

(ECA1-ECA6) to obtain a knowledge of the variation in PMD by time. The winding tension for

all fibres were approximately 30 grams. The results are shown in Figure 1 , ECA7 (representing

the wavelength scanning technique) and TELIA (representing the interferometric technique)

140
were measurements carried out at the same time. The results show that ECA representing the

wavelength scanning technique tends to measure a higher PMD-coefficient at higher PMD

values than TELIA representing the interferometric technique does.

Cabled fibres

Two 36-fibres ribbon cables were wound on cable drums with a diameter of 950 mm, the

winding tension was unknown but low enough to not impact on the results. Since the results are similar in both cables only PMD-coefficients from one of them are shown in Figure 2. The

PMD-coefficients ranging from 0.03 to 0.36 ps/Vkm. This gives a quite broad PMD range and

thereby a good indication of the relationship between the two methods. A statistical treatment,

using the t-distribution gives the following results, averaged over all fibres.

Interferometric Method: PMD=0.12±0.16 ps/Vkm (95% significance level)

Wavelength scanning method: PMD=0.13±0.16 ps/Vkm (95% significance level)

This shows a good correlation between measurements performed with both techniques.

0 350

Ribbon 101 201 102 103 203 104 105 205 106

Figure 2 PMD measured on a cable, the high PMD on 3 of the ribbons is due to the cable design.

141
Summary

In Figure 3 a plot of ECA-values against TELIA-values is shown, PMD-coefficients measured in both cables and on the spools with different diameters are taken into consideration. In the plot a linear regression is also shown, giving a con'elation coefficient, R, of 0.92. From a statistical

point of view a R-value > 0.9 is indicating a reliable relationship. Further, a statistical treatment

of the results by means of a two sided t-test with a significance level of 95% showing a reliable

dependence between the two methods. Results from field measurements in Sweden

are discussed in (4).

•

•

« ^

4

«

<

«

0.000 0.050 0.100 0.150 0.200 0.250 0.300 0.350 0.400 TELIA, the interfero metric techntque

Figure 3

Acknowledgement

We thank Dr A. Barlow from EG&G for valuable technical discussions.

References

1 N. Gisin, "Polarization Mode Dispersion:Definitions, Measurements and Statistics",

Symposium on Optical Fiber Measurements, 1994.

2 B. Arvidsson, H. Mickelsson, A. Barlow, "A PMD Study on Fibres in a Cable

Factory", OFMC 95.

3 FOTP-113, TIA.

4 K. Brising, "Polarisation Mode Dispersion Field Measurements: Survey of the Swedish

National Optical Network", to be published in IWCS 1996.

142
Interferometric Polarization Mode Dispersion Measurements with

Femtoseconds Sensitivity

Ph. Oberson<^, K. JuUiard^^, N. Gisin^,

*^Group of Applied Physics

University of Geneva

1211 Geneva 4

Switzerland

*Center for Telecommunications Studies

Pontifica Universidade Catolica

Rio de Janeiro 22453

Brasil

R. Passy*, JP Von der Weid*

The standard interferometric measurement technique for Polarization Mode Dispersion

(PMD) measures the cross correlation of the electro-magnetic field at the output of the fiber or

optical device under test when a broad spectrum light source is used [1]. The sensitivity

(smallest measurable PMD delay) of this well-established technique is limited by the

coherence time of the light source. In this contribution we analyze critically extensions of the

interferometric technique to lower values and present a new extension which provides a

sensitivity of a few femtoseconds.

When the polarization mode delay is smaller or comparable to the coherence time of

the source, the Fiber Under Test (PUT) behaves as a two mode fibers: the so-called principal

states of polarization. For example, if the Differential Group Delay (DGD) is smaller than 0.1

ps, then the FUT behaves as a two mode fiber even when illuminated with an LED with 80

nm spectral width. In this case, in order to measure the DGD between these two principal

states with the interferometric technique, the acquisition of the interferograms as well as their

analysis and interpretation have to be changed.

A first possibility consists in using a birefringent Michelson interferometer based on a

Polarization Beam Splitter (PBS) instead of the usual 50-50% beam splitter. The axes of this

PBS are then adjusted to the "axes" of the FUT. In this way the two arms of the

interferometers are extended by the two principal modes of the fiber. Hence the center of the

interferogram is shifted with respect to its position in the absence of the FUT. This method is related to the cross-correlation interferometer describe in [2]. Note however that this method

requires a full polarization controller to adjust the linear principal states of the birefringent

Michelson interferometer to the principal states of the fiber under test. Indeed, the latter are generally elliptic. Hence, a simple polarization rotator (like the half-wave plate used in [2]) is insufficient. With such a half-wave plate the measured DGD underestimates the real DGD.

Below we present for the first time a new interferometric high sensitivity measurement

technique based on a standard fiber interferometer with a 50-50% optical coupler (Patent

applied for). In this technique a full polarization controller plus a polarization maintaining

HiBi fiber of around 0.5 ps PMD delay are placed between the fiber under test and the

interferometer, see Figure 1. The HiBi fiber produces two side-peaks on the interferogram, left and right from the central aurocorrelation peak. When the principal states of the FUT are aligned with the eigenmodes of the HiBi fiber, their delays add, whereas they subtract when

the fast principal mode of the FUT is aligned to the slow mode of the HiBi fiber. Hence, by

143
measuring the displacement of the side-peaks when the orientation of the principal modes of

the FUT are scanned relative to the axes of the HiBi fiber, the delay between them can be

determined.

This measurement method is reminiscent of that presented in [3], since in both

methods a HiBi fiber is inserted before the interferogram. An important difference, however,

is that this reference [3] proposes to measure the spread of the interferogram (actually in [3]

the interferogram is computed by Fourier transform a wavelength spectrum, but this does not

change the argument). However, the determination of the position of an interferogram is much

more accurate than the determination of its width. Indeed the position can be determined with

a reproducibility of less than 3 fs [4], whereas it is difficult to determine the width with a

reproducibility better than 50 fs. For example, in [4] a standard deviation of 2. 14 fs was

obtained for a thousand measurements of the position of interferograms at 1.55 micron.

Moreover, without the polarization controller, the width of the satellite peaks do not contain

the full information about the DGD of the FUT. For instance, if the principal states of the

HiBi fiber and the FUT happen to be aligned, then the width of the satellite peak is determined

only by the LED spectrum.

According to our experimental results the sensitivity of our setup is of a few

femtoseconds (hence much lower than the source coherence time). The main limitation of this

new measurement method lies in the birefringence of the pigtails and other components of the

set-up. Note for example that three 1/4 wave plates can already produce a DGD of about 4 fs. Figure 2 presents three interferograms (around the position of one of the satellite peaks). They

represent the three different alignment conditions between the FUT and the HiBi fiber:

aligned, at 90 degrees and at 45 degrees, respectively. Clearly, the latter case corresponds to the larger width, as it should. The difference in position between the two extreme cases correspond to twice the DGD of the FUT. This result is in good agreement with other

measurement methods. Figure 3 presents two complete interferograms, showing the central

autocorrelation peak and the two satellite peaks for the two extreme alignments of the

principal states. The 1 1 fs result illustrate the high resolution of this technique. In Figure 4 a graph of measurement results for various length of a HiBi fiber of known PMD (0.23 ps/m) is presented. The horizontal scale represents the expected DGD directly proportional to the fiber

length. The vertical scale is the standard deviation of the distance between the two satellite

peaks over 125 interferograms obtains for different alignment conditions.

In conclusion, the existing extensions of the interferometric measurement technique

for PMD measurements to value lower than 0. 1 ps are of limited use. A new extension has

been developed and tested. It extends the measurement range down to about 10 fs.

Acknowledgments

This work has profited from stimulating discussions with Dr B. Ferny from the Swiss

Telecom PTT.

References

1. N. Gisin et al, IEEE J. Lightwave Technology 9, 821-827, 1991.

2. Y. Namihira et al. IEEE Photonics Technol. Lett. 1,329-331, 1989; see also OFMC'93,

Torino, p. 189-192, 1993.

3. A.J. Barlow et al., Proceedings of the 3rd Optical Fiber Measurement Conference, Liege,

Belgium, Edited by the European Action COST 241, p. II.4, 1995.

4. L. Thevenaz et al, IEEE J. Lightwave Technology 6, 1-7, 1988.

144
Figure 1 : Schematic of the set-up. The ellipse on the Fiber Under Test (FUT) indicates that

the principal polarization state are in general elliptic. The arrow on the HiBi fiber

indicate the axes of such fiber that define linear polarization states. The polarization

controller couples the two elliptic principal states of the FUT to the two linear

eigenmodes of the HiBi fiber. The remaining of the set-up is identical to the instruments

used for interferometric PMD measurements.

2.

I—

>

60

50

40

30

20

0.40

-30

-0,27 -0.13

At [ps]

0.00 0.13 0.27 0.40

'

1

'

' 2*50 fs

1 ' 1. ' 1

'

1 r-<i^C^-''\'.

1

(

l' 1 -20 -10 0 10 20 30

AL [|im]

Figure 2: Example of interferograms obtained for a FUT with a relatively large PMD of 50

fs. Only the scan around one the side peaks produced by the HiBi fiber is shown. The

full line corresponds to the case were the polarization controller is not aligned. The two

side peaks can still be distinguished. The dotted and dashed lines correspond to the case were the fast principal polarization states of the FUT are ahgned with the fast, slow

eigenmode of the HiBi fiber, respectively.

145
[ps]

-0.80 -0.67 -0.53 -0.40 -0.27 -0.13 0.00 0.13 0.27 0.40 0.53 0.67 0.80

->—I

'~ T T

2 * 11 fs

10 0 10

Al [i-im]

60

Figure 3: Example of interferograms obtained for a FUT with a DGD of about 10 fs. The

scans through the central autocorrelation peak and the side peaks produced by the HiBi

fiber are shown. The full and dashed lines correspond to the case were the fast principal

polarization states of the FUT are aligned with the fast, slow eigenmode of the HiBi fiber,

respectively.

60 i

1 1 1 1 r

r

0 - I I I 1 1 I I z 0 10 20 30 40 50 60 70 80

,

, Differencial groupe delay [fs]

Figure 4: Measurement results versus expected PMD. The horizontal scale correspond to 230

fs/m times the fiber length, where the PMD of 230 fs/m was measured on a longer fiber

sample with the standard interferometric measurement technique. The vertical scale

corresponds to one standard deviation of the measured distance between the side peaks in

125 interferograms obtained for different relative orientations of the FUT and HiBi fiber

axes. The standard deviation of the linear fit (full line) is of 5.5 fs. This illustrates the

sen.sitivity of the measurement technique and its linearity.

146
A Novel PMD Estimation Technique with

Wavelength Stepsize Optimization

Osamu Aso and Haruki Ogoshi

Opto-Technology Laboratory, Fumkawa Electric Co., Ltd.

6, Yawata-Kaigandori, Ichihara, Chiha, 290 Japan

Tel. +81-436-42-1724, Fax. +81-436-42-9340

e-mail: aso@optec.ch.furukawa.co.jp

1. Introduction

The concept of the principal states of polarization( PSP ) has been widely accepted and utilized in

polarization analysis. PSP was introduced by Poole and Wagner[l] as the stable state of polarization(

SOP ) in the frequency( wavelength ) domain.

In order to apply PSP theory to PMD analysis, the differential group delay( DGD ) time of the two

orthogonal PSPs is known to play an important role. According to Poole and Wagner, in the absence

of polarization dependent losses, Jones matrix U of the optical transmission medium is represented as

an element of the unimodular unitary matrices[2]:

U\{(x>) ujico)

U = ,detf/ = +l

mJC w) M*( w) J (!)

DGD time ^Ir of the PSPs is calculated as ^-^VISMif (2,

At present, in order to estimate the above At experimentally, two measurement techniques have been

developed:Jones matrix eigenanalysis[3] and Poincare sphere method[4],[5]. With the Poincare

sphere method, 3-dimensional real vector Q. is the essential parameter used to describe the

polarization properties of the transmission medium. The vector is called as the PMD vector and its

magnitude coincides with the DGD time of the PSPs:

According to the geometrical interpretation, the Jones vector is understood as the Euclidian spinor

quantity in Stokes space[2]. The transformation from the spinor field to the vector field makes it

possible to prove the mathematical relation of eq.(3) explicitly[6]. In this report, based on the result

of the above transformation, we will present a novel and direct PMD vector estimation technique

through the Jones matrix measurements.

2. Wavelength Dependence of the SOP in the Stokes and Jones Vector Space

As discussed in the preceding section, PSP is defined as the stable state in the wavelength domain.

147
In this section, we will brietly review the basic equations which describe the wavelength dependence

of the SOP both in the Stokes and Jones vector space.

2.1 Stokes Vector Representation

Stokes parameter representation of the wavelength dependence of a SOP has been discussed in

ref.[5]. Discussions are summarized by the following equation of rotation[4],[5]:

d s do) (4)

where s is the normalized Stokes vector and O) is the angular frequency of light. In eq.(4), Q

represents the rotational axis of the Stokes vector. The direction of the vector indicates the PSP of

the transmission medium, and the magnitude of the vector coincides with the DGD time of the

PSPs[5].

2.2 Jones Vector Representation

Jones vector is understood as the spinor representation of the normalized Stokes vector field[2]. Thus

any SOP points on the Poincare sphere are mapped into the spinor field. According to spinor

analysis, eq.(4) can be represented as[6],[7]

dco 2 ^,

where Wis the 2-dimensional complex Jones vector and H is the hermite matrix defined as

(5)

i do) -hi

hi

-h, (6)

9 I du '\ du^

h\ = -iMi -r-^ + U2

do) dcj

1 2 i du

hi = -\U2

do)

du2

dcj (J)

Using the unimodular unitary condition of the matrix U, it is easily shown that h\ is the real and hi is

the complex function. The physical meanings of eq.(4) and eq.(5) are essentially equivalent and both

represents the rotational motion of the SOP around the PSP[6].

3. PMD Vector Estimation Based on tlie Jones Matrix Measurements

As shown in eq.(4) and (5), wavelength dependence of the SOP is described in the similar manner

with rotation. The PMD vector is obtained from Stokes vector trajectory measurements with slight

wavelength changes[4],[5].

Based on the spinor transformation discussed in section 2. the PMD vector can be derived as:

h^ Im(/72)

LRe(/72)J (8)

The magnitude of the Q. vector is easily calculated and satisfies eq.(3). Thus, direct estimation of the

148
PMD vector is possible to realize through the Jones matrix measurement[3].

In order to be able to apply the above method, we must assess the suitability of the frequency step

size Ao) . If the step size is larger than the upper limit from the first order approximation, the higher

order effect[7] makes it difficult to estimate the PMD value accurately[5],[6]. For our discussions, a

suitable Ao) ensures the hermiticity of the matrix H: the element h\ must not have an imaginary

component. It is expected that the estimation of the following quantity will lead us to a suitable stepsize estimation

If ris not negligible, the first order approximation is violated and the estimated Q does not generally

represent the PMD vector of the medium. So a suitable threshold Fth (rth < 1 ) condition enables us to

ensure accurate estimation.

We will now provide a summary of the discussions of this section. Jones matrix measurement by the

sufficiently coherent light will enables us to determine the Jones matrix U(Ci)o) and U((i)i>+Aa)). Then

one can apply finite differential approximation in order to estimate the matrix H and the quantity F.

4. Numerical Simulation

In order to confirm the discussions in the preceding section, we will perform a numerical simulation

in this section. In order to show the availability of the scheme, we consider a transmission medium

with polarization mode coupling. The medium consists of two PMFs with finite angular

misalignment[9]. The specification of the two PMFs are as follows: they have the same linear

birefringence and that birefringence is assumed to be independent of wavelength. The DGD of each

fiber is set to be 3.00 ps and 2.00 ps respectively. Angular misalignment is set to be 30 deg..

Result of the simulation is shown in Fig. 1 . It is understood that when the step-size Aco decreases, I Q I converges to the theoretical DGD time 4.36 ps. From this result, we can say that measurement

feasibility is insured when we set Ft h = 0.10 with accuracy 0.07ps.

5. Conclusion and Remarks

We have developed a direct PMD vector estimation technique. However, the technique strongly

depends on the unitarity of the Jones matrix. On the other hand, at present, the practical Jones matrix

measurement procedure is not expected to preserve the geometrical symmetry of the matrix[3]. As the

exponential approximation scheme was introduced into the Jones matrix eigenanalysis by Heffner, it

is necessary to consider some approximation scheme [6].

(9)

149
1 0-4 1 0-3 10-2 10-1 1 10 1 02 1 03 r

Fig. 1 . Result of the numerical simulation. Horizontal axis shows

rvalue and vertical axis is the magnitude of PMD vector. Acknowledgment

The authors would Hke to thank Prof. T. Futamase and Mr. N. Jin of Tohoku University for their

helpful discussions. O. A. would like to thank Mr. H. Nakamura for his reading of the manuscript

and valuable comments. He also wishes to thank Mr. S. Yoon from the University of British

Columbia for his reading of the manuscript and continuous encouragements.

References

[1] C. D. Poole and R. E. Wagner, Electron. Lett., vol. 22, No. 19 ( 1986 ) pp. 1029 - 1030.

[2] H. Takenaka, A^OMv. Rev. Optique, vol. 4, No. 1 ( 1973 ) pp. 37-41.

[3] B. L. Heffner, Photon. Technol. Lett., vol. 4, No. 9 ( 1992 ) pp. 1066-1069.

[4]Y. Namihira and J. Maeda, Symp. on Optical Fiber Measurements, Boulder, Co. USA ( 1992 )

pp. 145-150.

[5] C. D. Poole, N. S. Bergano, R. E. Wagner and H. J. Shulte, J. Lightwave Technol., vol. 6, No.

7 (1988 ) pp. 11 85-1 190, and references therein.

[6] O. Aso, I. Ohshima and H. Ogoshi, " Feasibility Study on A Jones Matrix Measurement Method

Preserving A Geometrical Symmetry and Its Application to Polarization Mode Dispersion Analysis ",

to be submitted.

[7]C. D. Poole and C. R. Giles, Opt. Lett., vol. 13, No. 2 ( 1988 ) pp. 155-157.

[8] O. Aso, I. Ohshima and H. Ogoshi, OEC'94, Makuhari, Japan ( 1994 ) paper 14E - 2, pp. 238 -

239;Symp. on Optica]Fiber Measurement, Boulder Co, USA ( 1994 ) pp. 159 - 160.

[9] O. Aso and H. Nakamura, Electron. Lett., vol. 32, No. 6 ( 1996 ) pp. 578 - 579.

150
First Report of ITU-T PMD Round Robin Measurements

for Optical Fibers and Component

Y. Namihira" and ITU-T PMD Round Robin Group

[Asia (Japan)t'*'S Europe''''^ USA^^'^]

[*] KDD R&D Laboratories (ITU-T Coordinator)

2-1-15, Ohara, Kamifukuoka, Saitama 356 Japan

Tel : -^8 1-492-78-7801, Fax : +81-492-63-9328, E-mail: namihira@lab.kdd.co.jp

M.Ohashi (NTT), Y. Suetsugu (Sumitomo), M. Sawada (Fujikuia), O. Aso (Furukawa),

S. Suzuki (Hitachi Cable), M. Noguchi (OCC), M. Ichimura (Santec),

H. Fujishima (Oyokoden LAB), T. Ozeki (Sophia University)

[#2] B. Pemy (Swiss PTT, COST-241 Coordinator), N. Gisin (GAP), A. J. Barlow (EG&G)

[#3] P. A. WilUams (NIST, TIA Coordinator), K. S. Kim (AT&T), A. F. Judy (Lucent)

1. Introduction

In high bit rate, long haul, WDM and analog transmission optical amplifier systems, the optical

pulse broadening due to polarization mode dspersion (PMD) could impair the performance of the systans.

Therefore, it is voy important to measure accurately the PMD of the optical fibers and optical

components.

At the ITU-T WP-4/15 (Optical fiber cables and Systems) meeting in May 1994, they have been

decided to start the international PMD round robin measurements. It was agreed Uiat the three mnning

round robin measuronoits, one in Asia (Japan), organized by KDD (Y. Namihira), one in Europe,

organized by COST 241 (B. Pemy, Swiss Telecom PTT), and a third one in North America, organized by

TIA (P. A. Williams, NIST) would be combined in the following ITU-T meeting. Recendy, the PMD

round robin result's reports were presented at ITU-T SG15 May/June 1996 meeting [1]. This paper is die first report of ITU-T PMD round robin measuronents for single mode optical fibers and fiber optic component in Asian (Japan) region using KDD PMD specimens. The results of an

extensive inta-comparison of the ITU-T PMD round robin measurements on sevai kinds of KDD PMD

specimais, coordnated by KDD (Y. Namihira) are presented and discussed Sixteai laboratories from all over the world participated die intercomparison of the PMD round robin measurements.

2. Classification and Definitions of PMD Measurement Methods

The PMD is die cfifferential group delay (or PMD) Ax between two orthogonal state of polarization

( SOP). In the PMD measurements, sevaal methods have beai reported [2]. So fer, the PMD has been

measured by mainly two cifferent procedures. One is the time domain measurement and the odier is the frequency domain measurement. The former method is based on the Interferometric method {Air path -

Crosscorrelation type (Int-X) [2,3] and Fiber refaence-Autocorrelation type (Int-A) [4] } and Optical pulse

method [2]. The latto" method is based on the evolution of the states of polarization (SOP) as a function

of the fi^uency (wavelength).

In the frequaicy domain, diere are six methods (wavelength scanning method) such as the Fixed

analyzer method with extrema counting (FAEC) mediod [5,6], the Fixed analyzer method with Fourier

transfonnation analysis (FAFT) mediod [7], die Poincar^ sphae (PS) mediod [2,8-10], Uie SOP mediod

[2], die Jones matrix eigenanalysis (JME) method [11] and the Transfa function matrix (TFM) method

[12]. The classification of the various PMD measurement mediods are sunmiarized in Table 1

3. PMD Round Robin Measurements

Three block's ITU-T PMD round robin measurements coordinators of Asia (Japan), Europe (COST

241) and USA (TIA) are shown in Fig.l.

The photograph and specifications of KDD PMD specimens are shown in Fig.2 and Table 2, respectively. This KDD PMD specimens are still available. The fu-st results of ITU-T PMD round robin measurements of die KDD specimens KDD-#1, #2, #3,

151
#4, #5, #6 and #7 are summarized in Table 3. The dimensionless values (normalized PMD) in the Table 3

are experimentally measured values normalized to the average value measured by the all participants.

No (weak) mode coupling case of KDD-#1, #2 and #3, strong mode coupling case of #4 and #5, and

weak (strong) mode coupling case #6 and #7 are shown in Fig. 3 (a), (b) and (c), respectively.

From Table 3 and Figs. 3 (a), (b) and (c), it was found that

1 ) Interferomeuic method using crosscorrelation definition (Int-X) [2,3] and Jones matrix eigenanalysis

method (JME) [11] showed consistent results comparing with the other methods.

2 ) Interferometric method using autocorrelation definition (Int-A) [3] showed unstable results in the case of no(weak) (KDD-#l-3) and weak(strong) (#6,7) mode coupling cases, but in the strong mode coupl

-ing case, stable results were obtained. Especially, one results of #1 showed extremely larger values. 3 ) Fixed analyzer method using extrema counting (FAEC) [5,6] showed a slightly larger values comparing with the Int-X and JME methods in the strong mode coupling case. Accordingly, in the strong

mode coupling sample, FAEC method may require the mode coupling correction factor (k < 1). 4 ) FA method using Fourier transformation analysis (FAFT) [7] showed larger measurement results of

KDD-#5 and #6 comparing with the other methods (Int-X and JME).

5) The PMD values of the Poincare sphere (PS) [2,8-10] and SOP [2] methods were obtained from the same Stokes analyzer (Polarimeter), the both calculating algorisms were different The both results were almost the same. But, the SOP method's measurement dynamic range was narrower than the other methods.

6 ) The PMD values of the JME [11] and Transfer function matrix (TFM) [12] methods were obtained

from the same Stokes analyzer, the both calculatring algorisms were different. The TFM measurement

results showed extremely smaller measurement results of #1 and #7 comparing with the other methods.

5. Conclusion

It was confirmed that the Intoferometric method using crosscorrelation definition (Lit-X), Poincare' sphere (PS) and Jones matrix eigenanalysis (JME) methods showed consistent results comparing with the other methods such as Autocorrelation type Interferometric method (Int-A), Fixed analyzer using extrema

counting method (EAEC), Fixed analyzer using Fourier transformation analysis mediod (FAFT), SOP

method and Transfer function matrix method (TFM), respectively.

References:

[I] COM 15-301 (KDD): "First report of ITU-T international PMD round robin measurements in Asia

(Japan), Europe and USA", ITU-T SG15, May/June, Geneva, 1996.

[2] Y. Namihira et al.,: "Polarization mode dispersion measurements in optical fibers", (Invited papers),

Symposium on Optical Fiber Measurements (SOFM'92), NIST, Boulder, Co., USA, p.l45, 1992.

[3] Y. Namihira et al.,: "Comparison of various polarisation mode dispersion measurement methods in opUcal fibres". Electron. Lett., vol.28, no.25, p.2265, 1992.

[4] N. Gisin et al., :"Polarization mode dispersion of short and long single-mode fibers", IEEE,

JLT,vol.9, no.7, pp.821-827, 1991.

[5] Y. Namihira et al.,:"PMD reduction of optical fiber cables for transoceanic optical amplifier

submarine cable systems" (The Best Paper Award), rWCS'93, Session 14, pp.655-664, USA, 1993.

[6] C. D. Poole et al., : "Polarization -mode dispersion measurements based on transmission spectra

through a polarizer", IEEE,JLT,vol.l2, No.6, p.917, 1994.

[7] A. Barlow et al.,: "Polarization mode dispersion measurement using Fourier analysis: Investigation

of mode coupling behaviour in fibers", SOFM'94, pp.167-170, NIST, Boulder, Co., USA. 1994.

[8] Y. Namihira et al.,:"Polarisation mode dispersion measurements in 1520 km EDFA system".

Electron. Lett., vol.28, no.9, p.881-882, 1992.

[9] C. D. Poole et al., :" Polarization dispersion and principal states in a 147-km undersea lightwave

cable", IEEE, JILT, vol.6, No.7, pp.1185-1189, 1988.

[10] D. Andreciani et al., ;"Measurement of the gioup-delay difference between the principal states of thp

polarization on a low-birefringence terrestrial fiber cable". Opt. Lett., vol.12, no.lO, p. 844, 1987.

[II] B. L. Hef&ier : "Automated measurement of polarization mode dispersion using Jones matrix

eigenanalysis", Photonics Technol. Lett., vol.4, no.9, p. 1066, 1992.

[12] T. Ozeki et al. :"Birefiingence distribution along fiber length", OFC'96, FA4, p.295, USA, 1996.

152
Europe Asia (Japan)

KDD

(ITU-T)

USA

NIST

(TU)

Dr. Y. Namlhlra

(*) Coordinator (ITU-T)

Dr. B. Pemy Dr. P. A. WUUams

Fig. 1 Coordinators of ITU-T International PMD round robin measurements

Table 1 Classification and definitions of various PMD measurement methods

Time Domain Frequency Domain (Wavelength Scanning)

Optical Pulse

Power Principal State of Polarization (PSP)

HI lerieruu Fixed Analyzer (FA) Jones Matrix Eigenanalysis(JME)

Poincare' Sphere

(PS)

State of

Polarization (SOP)

Transfer Function Matrix (TFM)

CrosscorreUtion

(Air Path Type)

ant-X)

Autocorrelation

(Fibre Ref. Type

(Int-A)

Extrema

Counting (FAEC)

Fourier Transformation Analysis (FAFT)

Optical Source

Broadband

Source

(BS)

Broadband

Source

(BS) Pulsed-LD (#1)BS

{#2)Tunable-LD

(T-LD)

Broadband

Source

(BS)

Tunable-LD

(T-U)) Tunable-LD

(T-LD)

(#l)T-LD

(#2) BS & MON Tunable-LD

(T-LD)

Analyzer

or Receiver

Michelson (or Mach Zehender)

Interferometer

Michelson (or Mach Zehender)

Interferometer

Streak Caooera

(#1) Optical Spectrum

Analyzer (OSA)

(#2) Optical Power Meter (OPM)

Optical Spectram

Analyser (OSA)

Stokes Analyzer

Stokes Analyzer

(#l)Sloke<! Analyzer

(#2) Rotatable Analyzer or Polarimeter

Stokes Analyzer

Basic Physical Measurement

Crosse orrclatioa between two

ortbogoi]al polarisation modes

AutocorrelaiioQ between two

orthogonal

polarization modes

A separation between two

orthogonal

polarization pulse

Transmission

through an

output polarizer

at a aeries of

wavclegnths

Transmission

tbrough an

output polarizei

at a series of

optical frequency

Jones matrix

at a scries of wavelengths

Stokes vector

arc on Pbincare

sphere

at a series of wavelengths

SOP curve

a5 a function of wavelength

Transfer

fii action matrix

at a series of wavelengths

Definition of PMD

(or DGD)

(#1) A half Peak

separation of two inlerfero- graiDs [Weak]

{1*2) Square root of the second mOCDCDt of

inlerfcrogram

[Strong]

Square root of the second ToowcnL of

interferogram

Peak

separation of two

optical pulses

Derive from number of extrcma and

wavelengths

of first and

lastextremum

Square root of the second mocoent of

Fourier transformed wavelength

scanning

data

Average of the DGD values computed

from JoDes

matrix pairs

Derive from

Poincare

sphere rotatic«i angle and

frequency

interval

Derive from a half number of extrcma and

wavelengths of

first and last extrcmum of SOP

curve

Average of the PMD valiKs computed from

Machraurin

exapansion using Transfer matrix

Fig. 2 Photograph of KDD PMD specimens

Table 2 Specifications of KDD PMD specimens

KDD PMD Test Specimens Length

[m]

Loss

[dB] PMD

[psl Mode

Coupling Model Contents KDD-#1 Optical Isolator -0.04 -O.S < 0.1 Weak

KDD-#2 No Spliced PMF -0.646 -3.2 -0.1 No

KDD-#3 No Spliced PMF -0.67 - 1.3 - 1.0 No (Weak)

KDD-#4 SOx PMFs

(Kandomly spliced) - 10.0 -9.9 -2.0 Strong

KDD-#5 6x PMFs

(Kandomly spliced) -8.708 - 1.8 < 5.0 Strong

KDD-#6 G.654 Fiber Spool (CSF) - 13^91 -3.0 < 0.1 Weak

KDD-#7

G.653 Fiber Spool (DSF)

- 13,673 -3.7 < 0.2 Weak

(Strong) PMF : Polarization Maintaining Fiber (Hi-Bi Fiber) CSF tCut-ofr Shifted Fiber aTU-T G.654) DSF : Dispersion Shifted Fiber (ITU-T G.653)

153
Table 3 Results of ITU-T PMD round robin measurements for optical fibers and component using KDD PMD specimens [The dimensionless values

(normalized PMD) in the Table 3 are experimentally measured values normalized

to the average value measured by the all participants]

PMD Method KDD#I 1 KDDi2 KDD»3 KDM4 KDCMS KDD** 1 KDD#7 Equipment Alxariim

Inl-X (A) 0.7S9 l.OM « » # « CKO CKO

Inl-X (R) 1.014 l.OOS I.IU l.MI 0.751 0.941 CUfl CUO

Inl-A (R) i.in 0.»7» }.«9« I.IOl 1.07} 1.40} l.t}9 CKD^l) C}(1)^1) FAEC (M) \.m l.OM 1.00} 1.034 0X33 l.KS 1.111 C4f7)^(D,C(S(4) C4(T),CS(D^t(fl FAFT(R) \.t\i I.OIS 0.991 1.175 IJ«4 lJU 0.9iS C7(2) JME{M) S.9g9 1.023 l.M« *JU OJiS tMf *M1 C»(t1 JME(R) ».n9 l.OU l.OOl •.9S( 0.99} 0.1S9 0.9M C4(0 CS«) PS(M) 0.9a* t.eo* 0.999 •.9*7 0.933 •J49 t.m P1(1)J*}(1) SOP(M) *

l.ftS3 0.994 I.MS # #l.OJS P1(l) pi(i) TFM (M) o.cis 1.15J 1.0O4 •.M9 •J13 0.790 0.411 C«(l) Hl(I) Mode Coupling Wt.k No SIronc Wt.k(Stnnf) W«.k(SlMnf) Bquipm.nl

(f) Not Meajimd C i CommcrdaJ

Inl-X(A) ; Inlerferometric CroMeorTrlilioB AiT-p«ih typo (Pe»k) Int-X(R) ; Into-fcrometric Crojicorrelalion Air-p«th type (RMS)

Int-A(R) ; Inlerferometric Aulocorrciition Fiber reference type (RMS) FAEC (M) ; Fixed Analyzer Enremi Counting. [Mein] FAFT (R) ; Fixed Analyzer Fourier Ation An«lyii». [RMS]

1.00

Pi Prolo<j>^ H I HomcmaJd { ) [ Number

JME(M) ; Jone* Matrix Eigeoanalyiia (Mean) ;ME(R) ; Jona Matrix Eigenanalyaia (RMS)

PS(M) ; Poincare Sphere (Mean) SOP(M) ; Sute of Polarization (Mean) TFM (M) ; Tnnafer Function Matrix (Mean)

0.

OZ

1.73

IJO

0.73

OJO

0.23

\ No (veai) Mode CoapUng

(•) / \ (@KDD PMD Sptdmtiu)

\ 01. n, 13 -•-EDOO

IH-X(A) iB.Xfll) IH-AfK) FABC(»0 PAPT (S) !MZ(Kt) afflfS) PS (M) 30P(M) TTM (M) PMD MeAo4t

2.» r- -

1.73

OZ

(b) Strong Mode CoapUng

ConcaUned PMFt

(<»KDD PMD Spednuns ) *4, tS

-ODM

023

I«.X(A) la X(Tl) I« A(1» PAEC(M) PAFTfS) IME (M) n><E(B) PS fM) SOP (M) tTM (Ml PMD MtHwds

2.00 p - -

a.

oZ

(C)

Wfoil (Strong) Mode Coupling Long Optical Flien (9KDD PMD Soedment)

t6,ii

I».X(A) 1K-X(») IK AW PAEC(M) PAPT(S) n4E(*0 n<Bfli) PS 00 SOP (M) TFM (M) PMD Mtthodx

Fig.4 Results of ITU-T PMD round robin measurements (Normalized PMD)

@ KDD PMD Specimens

(a) No (weak) mode coupling; Optical isolator & No spliced PMF

(b) Strong mode coupling ; Ramdoraly spliced PMFs

(c) Weak (strong) mode coupling; Long optical fibers

154
TIA round robin for the measurement of PMD

P.A. Williams

Optoelectronics Division - National Institute ofStandards and Technology

325 Broadway - Boulder CO 80303, U.S.A.

From 1994 to 1996, the National Institute of Standards and Technology (NIST) conducted a round

robin for the measurement ofpolarization mode dispersion (PMD). This comparison was coordinated

through the Telecommunications Industry Association (TIA) and involved participants from the

United States as well as Europe and Japan measuring three PMD specimens. This paper will report

the results of that round robin.

Three specimens were circulated to test various aspects of the measurement. These specimens were: Specimen 1, a pigtailed quartz waveplate with no polarization mode coupling; Specimen 2, a modecoupled artifact composed of a stack of 12 quartz waveplates cemented together and pigtailed; and

finally, Specimen 3, a 15 cm diameter spool of 25 km of typical communication fi.ber. All three

specimens had a nominal PMD of a few tenths of one picosecond. This round robin included 1

participants from the United States, Europe, and Japan, and used up to four different measurement

techniques: Jones matrix eigenanalysis (JME) - or equivalently the Stokes method, interferometry,

and wavelength scanning, also known as fixed analyzer, with either extremum counting (WSEC) or Fourier transform analysis (WSFT) [1].

Specimen 1 (Stable, non-mode-coupled)

The single quartz waveplate has a

theoretically predictable PMD which was

calculated from the thickness of the

waveplate and its wavelength dependent

birefringence. However, the 2-3 m long

fiber pigtails of the specimen have a nonnegligible PMD which dominates the

uncertainty of the device PMD. It is therefore possible to predict the measured

PMD as the wavelength-averaged

differential group delay DGD of the single

waveplate device only to within ±2.5%,

where the uncertainty is the maximum

PMD variation seen experimentally when

the fiber leads are arranged in various

orientations.

Figure 1 shows the variation among

participants' measurements of Specimen 1

The x-axis represents the participants, and

the y-axis is their measured PMD normalized to the theoretically predicted value for the wavelength

.a

1.3

1.2

1.1

1.0

— o

g 0.9

0.8

I I I I I I I I I I I I I I

f . JME

o WSEC

X WSFT

o Interferometer

o

• <> o o

Oo

0.7 ^

I I I I j. I I I I I I I

I I I I I I I I r 5

' ' lO' 15' 20

Participant Number

25

Figure 1 Single waveplate (Specimen 1).

155
at which the measurement was made. There are more participant numbers on the x-axis than

participants because some participants made measurements on more than one similar system and

consequently were assigned two participant numbers. Table I shows the measurement statistics. The

agreement among participants is good, with most of the variation explainable by the PMD

uncertainty due to the fiber leads (dashed lines of Figure 1). A few significant disagreements with

theory can be seen. Some of these disagreements are explained by improper measurement

techniques. One such disagreement is apparent in some of the interferometric results, where

measurements were made by calculating the second moment of the double-peaked interferogram

rather than measuring the time delay between the peaks. The existence of the autocorrelation peak

in the interferogram might have led to incorrect results (yielding the smaller than theoretical values).

Table I Average / standard deviation ofnormalized measurements for

each specimen. *One extreme data point was removed for

interferometry (Int) statistics.

Specimen 1 Specimen 2 Specimen 3

WSEC 1.00/0.067 1.02/0.119 1.08/0.28

WSFT 0.99/0.014 1.00/0.038 1.014/0.13

Int 0.94 / 0.068 0.99/0.051 0.96/0.21

*

JME 0.99 / 0.022 1.00/0.010 0.92 / 0.22

Total 0.98/0.060 1.00/0.076 n.a. / 0.24

Specimen 2 (Stable, mode-coupled)

The second specimen also has a theoretically predictable PMD. However, its application is not as

straightforward. The DGD ofthe waveplate stack is derived from the measured thickness and known

birefringence of each plate along with the known orientation angles of their optic axes with respect

to each other. From this information, a Jones matrix for the 12-plate stack was calculated and DGD

as a function of wavelength was derived. This information allows a theoretical prediction of PMD

(as the average DGD over wavelength) with an accuracy similar to that of the single waveplate

where the limit is a ± 2% uncertainty due to PMD of the fiber leads. This evaluation method allows

accurate theoretical prediction of JME measurement results.

There is, however, a complication for other measurement techniques. A theoretical prediction of

PMD is more difficult for techniques which do not measure DGD directly but rather measure PMD

through some statistical parameter. Such techniques as interferometry and wavelength scanning vsdth either extremum counting or Fourier analysis are subject to the fact that an exact theoretical

prediction of the PMD of the mode-coupled artifact is possible only if the orientations of the

measuring polarizer and analyzer are known with respect to the axes of the waveplate stack. This is not possible since the fiber pigtails randomly alter the polarization state before it impinges on the

waveplate stack. In order to establish a theoretical prediction, participants made between 3 and 6

156
measurements with the polarizer and analyzer

randomly rotated (or alternatively with the

lay of the fiber leads randomized) between

each measurement. The average of these

measurements was normalized to a

theoretical average derived from a computer

simulation of 100 measurements over the

same wavelength range using random

polarizer and analyzer orientations.

Results are shown in Figure 2 and Table I. As

expected, the best agreement came from the

JME measurements where all of the data

points fell within the uncertainty due to the

PMD of the fiber leads (dashed lines). For

the rest of the measurement techniques, most

ofthe disagreements between experiment and

theory are within the 2a/N'''' error bars, where a is the standard deviation of the theoretical

prediction and N is the number of

measurements made by the participant.

Uncertainties of the participant's

measurement system are not included.

Specimen 3 (Spooledfiber)

Finally, for comparison Figure 3 and Table

I show the results of measurements made on

the spool of 25 km of single-mode fiber. Participants' results are plotted as the

measured value divided by the average

measured value of all the participants. There

is no theoretical prediction with which to compare, but the standard deviation among

participants gives an indication of the

repeatability which can be expected on such

a spooled fiber.

Conclusion

For measurements on the fiber spool

(Specimen 3), the standard deviation o was 24% with no statistically significant

disagreements between measurement

techniques.

1.3

1.2

1.1

1.0

0.9

0.8

0.7

1—I—I—I—

I—I—1—1—I—I—

TO

-

TIME ^

o WSEC

X WSFI

o Interferometer

T—I—

I—I—I—I—I—I—I—I—I—

<> -- C)

J I I I I I I I I I I I I I I ' I I ' <^ I I I ' 0 5 10 15 210 25

Participant Number

Figure 2 Waveplate stack (Specimen 2).

2.0

1.5

I 1.0

0.5 -

0.0

"T I I I

I

I I I I I

I I I I

I

I I I 1

I

I I I r (TIME

o WSEC

X WSFT

o Interferometer o o

-0

• X

o • o

• • 8o

I I I 'I I I I I I I I I I I I I I I I '

I ' 5 10 15 20 25

Participant Number

Figure 3 Spooled 25 km fiber (Specimen 3).

The non-mode-coupled Specimen 1 shows significant improvement over the fiber spool. With the

157
exception of interferometry, all of the measurement techniques yielded averages among the

participants which were within 1% of theory. The overall uncertainty of specimen 1 is only 6%. The

WSFT and JME techniques performed best with sigmas of only 1.4% and 2.2%, respectively (the

WSFT result represents only 4 measurements). The poorer repeatability of WSEC and interferometry

might be due to applying mode coupled measuring techniques to measuring the non-mode-coupled

device.

For mode-coupled Specimen 2, each measurement technique's average value agrees within 2% of

theory. The theory predicts only what the participant will measure using a given technique. It makes

no statement as to whether different techniques should measure the same value. Indeed for this

moderately mode-coupled 12-waveplate stack, all the techniques are not expected to measure the

same value, but the normalized data does not illustrate these systematic measurement differences.

With regard to standard deviation, the overall o of 7.6% is only slightly higher than for Specimen

1 and is significantly lower than for the fiber spool. The JME technique shows the best stability

among participants, with a a of only 1%. This is as expected, since the JME technique is only

weakly affected by birefringence in the fiber leads. Based on the results of this round robin, NIST

is constructing a stack of 35 waveplates to increase mode-coupling and PMD. We anticipate that this

new artifact will reduce the uncertainty due to PMD of fiber leads for all of the measurement

techniques.

The author expresses thanks to the members of the TIA who helped in the organization of this round

robin and also to all the participants from the United States, Japan, Switzerland, and the United

Kingdom for taking part.

[1] For measurement methods using these techniques, see Fiber Optic Test Procedure (FOTP)

113, 122, and 124, Telecommunications Industry Association.

158
MEASUREMENT OF SOP EVOLUTION ALONG A LINEAR

BIREFRINGENT FIBRE WITH TWIST USING POLARISATION OTDR

RE Schuh and AS Siddiqui

Department of Electronic Systems Engineering,

University of Essex,

Colchester, C04 3SQ, UK

Abstract: We show how the state of polarisation (SOP) evolution along optical fibres can be measured using the non-destructive method of polarisation optical time domain

reflectometry (POTDR). From the SOP evolution the linear birefringence and twist can be calculated along the fibre, thus identifying fibre sections with and without twist

and sections with relatively high and low polarisation mode dispersion (PMD).

Introduction: POTDR, like OTDR, is a non-destructive method for measuring fibre

parameters such as the birefringence characteristics [1] along the fibre with access to only one end. If the linear and circular birefringence distribution along a fibre is known, the fibre can be modelled. This would allow not just the PMD of the fibre to be calculated [2] but, as we show in this paper, also fibre sections with and without

twist to be identified.

The POTDR fluctuation is a complicated function of the linear-circular birefringence

distribution along the fibre. We recently showed that fibre twist significantly

influences the periodicity of the observed signal from POTDR, and has to be taken

into account when estimating PMD using this technique [3, 4]. In this paper we report a further result, namely that the SOP evolution obtained from a POTDR can be used

to give the polarisation properties of the fibre. We believe this is the first report of

experimental POTDR results showing the SOP evolution on the Poincare sphere along

a fibre with twist.

Theoretical Model: In [3, 4] it was shown that S(/) , the SOP observed at the input

end of the fibre from the backscattered signal at the length /, along a uniformly twisted

fibre, can be expressed as S(0 = R^(P/)A^R(P/)S(0) = M(p/)S(0) (1)

where Am is the rotation invariant mirror matrix which changes the handedness of the

SOP and R^ is the transpose of a rotation R, given by

R = I-Fsin(P/)B-h(l-cos(p/))B' (2)

3 where I is the 3x3 unit matrix, B^ = X^y*P/t ^ skew-symmetric matrix where e^k is

k=\

the permutation symbol which is 1 for cyclic, - 1 for non-cyclic permutation of indices

1, 2, 3, and the unit vector |3 is determined in the rotating reference frame as

P = 0 2y-gyJ/^. Here, the magnitude, ^=^j^l +(2j - gyf is the

resultant birefringence due to the linear birefringence, Pl, the twist rate, y, and the

rotation coefficient, g. The occurrence of the rotation matrix R twice in Eq. 1 indicates

that the matrix coefficients of M are quadratic in cos(P/) and sin(P/) and in general the

intensity through an analyser may be expressed in the form of a Fourier series

159
I =aQ + ^a^cos{n^l) + b^^sin(n^l) (3)

n = l In the general case the evolution of the backscattered SOP is a self intersecting 'figure

of eight' whose detailed shape depends on the incident SOP, linear birefringence and

twist induced circular birefringence. For the simple case of zero twist the SOP trace is just a circle, ai and bi being zero, leading to just one component in the POTDR power

spectrum. For backreflection from a section with twist, ai and bi are now not zero which, together with ai and bi, lead to two components in the POTDR spectrum. This

qualitative difference between the two POTDR spectra makes it possible to detect

sections of fibre with twist. Measurement with simulation results using the power

spectrum with the two frequencies will be presented. In this summary we show the

full SOP evolution along the fibre obtained from POTDR.

Equation 1 may be extended to the SOP evolution as measured with the POTDR set- up shown in Fig. 1 by multiplying the backscattered SOP, S(/) , with a constant

rotation matrix, Rc which is the transmission matrix of the 3dB coupler in backward

direction, so that

S(/) = RcM((3/)S(0) (4)

Rc is just an overall rotation of the backscattered SOP trace on the Poincare sphere.

Results and Discussion: A commercially available OTDR, with a spatial resolution

of 2m, was modified with a computer controlled rotatable X/4 waveplate-linear

polariser combination. Fig. 1 . A 56m length of fibre was freely suspended from either

end, and twisted at the far end. The zero twist condition could be found by measuring

the PMD at different twist rates [4]. The SOP from the backscattered intensity can be

calculated by rotating the X/4 plate to at least four independent positions. In our experiment we used 19 analyser positions from 0 to 180 degree for calculating the

backscattered SOP, using a least square method. The SOP data was very reproducible over repeated measurements. Due to the dead zones at either end only 50 metres of the

fibre could be resolved with our POTDR setup.

Fig. 2(a-e) shows the measurement and best fit simulation results at different twist

rates for the backscattered SOP. The simulation is fitted to the measurements

interactively on both the Poincare sphere and on the azimuth and ellipticity variation

of the SOP as a direct function of length using Eq. 4. Figure 2(a) shows the SOP

evolution on the Poincare sphere for the fibre in the zero twist condition, defined as

that giving maximum PMD [4]. Now if the zero twist condition actually existed over

the whole length of the fibre then the SOP trace would be just a repeated circle.

However in the data of Fig. 2(a) only the last 3 1 metres of the fibre shows a repeated

circular trace on the Poincare sphere, the first 19 metres giving data points away from

the circle. From this, and measuring the fibre with left and right hand twist using the

POTDR, we could conclude that the first 19 metres of fibre has some small non- uniform linear birefringence and twist which may be frozen in during the fibre

fabrication process. The last 3 1 metres of fibre was then modelled as a section with a nominally constant linear birefringence as the variable, to find the best fit to the

measurements. The best fit simulation, indicated in Fig. 2(a), was obtained for |3l =

0.19rad/m for the last 31 metres of fibre. This value is within 34% of a value obtained

by a direct PMD measurement over the whole 56 metre length of the fibre.

160
In Fig. 2(b-e) we again carry out the simulation for the last 31 metre section of the

fibre. Applying a twist of 0.01 turns/m, Fig. 2(b), the backscattered SOP traces a

figure of eight shape in accordance with Eq. 2 and 4. For the higher applied twist rate of 0.019 tums/m, Fig. 2(c), the area of the figure of eight decreases as predicted in [3,

4]. The simulated traces for the two applied twist rates, Fig. 2(b-e), again give {3l =

0.19rad/m and twist values of 0.012 and 0.021 turns/m respectively for the best fit. Fig. 2(d-e) show the azimuth and ellipticity variation of the SOP evolution for the

applied twist of 0.019 tums/m as a function of length, where L = 27i/p is the input

SOP independent periodicity of the traces. Applying twist in the opposite direction

also gave a figure of eight in the SOP evolution but with opposite direction as expected.

Conclusions: The SOP evolution along a fibre has been measured using POTDR and

plotted on the Poincare sphere. It has been shown that the linear birefringence and

twist along the fibre can be evaluated using this data. Knowing the twist and linear

birefringence along the fibre the PMD of the fibre can be calculated. This may be used

during the manufacture of the fibre at different production stages to monitor the

intrinsic PMD expectation of the fibre.

References

[1] A. J. Rogers, "Polarization-optical time domain reflectometry: a technique for the measurement

of field distributions", Applied Optics, vol. 20, no. 6, pp 1060-1074, 1981

.

[2] R.E. Schuh, E.S.R. Sikora, N.G. Walker, A.S. Siddiqui, L.M. Gleeson, D.H.O. Bebbington,

"Theoretical analysis and measurement of the effects of fibre twist on the differential group delay

of optical fibres". Electron. Lett. vol. 31, no. 20, pp 1772-1773, 1995.

[3] R.E. Schuh, JG Ellison, L.M. Gleeson, E.S.R. Sikora, A.S. Siddiqui, N.G. Walker,

D.H.O. Bebbington, "Theoretical analysis and measurement of the effect of fibre twist on the polarisation OTDR of optical fibres", OFC'96, FAS, San Jose, 1996.

[4] R.E. Schuh, JG Ellison, AS Siddiqui, D.H.O. Bebbington, "Polarisation OTDR measurements

and theoretical analysis on fibres with twist and their implications for estimation of PMD",

Electron. Lett. vol. 32, no. 4, pp 387-388, 1996. Acknowledgments: The authors wish to acknowledge X. Shan and J.G. Ellison for helping with the measurements and D.H.O. Bebbington for helpful discussion. This work was carried out under EPSRC

Research Grant, number GR/K49454. The industrial partners are BT Laboratories, Martlesham Heath,

Ipswich and Alcatel Submarine Networks, Greenwich, London.

20 ns R(p/)

OTDR

1.55[im

JL 3dB

Collimators

Adjustable

Linear

Polariser

Adjustable

mplate

Fig. 1 Experimental POTDR setup.

161

Polarisation Dependent Loss of Concatenated

Passive Optical Components

A. ELAMARl\ N. GISIn\ B. PERNY^, H. ZBINDEN' , Ch. ZIMMER^

University of Geneva, GAP/0, 20, rue de I'ecole de medecine, CH -1211, Geneva 4

Swiss Telecom PTT, F&E 221, CH- 3000, Berne

Abstract

Global Polarisation dependent loss ( PDL ) of concatenated elements, each with some PDL, is difficult to predict since PDL can not be simply added. Moreover, if the fibre between the PDL elements has random effects on the polarisation state, then the global PDL is also random. We present simulations and measurements results of the corresponding statistics.

In Passive Optical Networks (PONs), telecommunication optical fibres are usually combined with optical components

like optical fibre couplers, wave length division multiplexers, isolators and other optical devices each of them may

have an amount of polarisation dependent loss. Hence it is necessary to evaluate the global cumulated PDL of the concatenated elements. Contrarily to what one could expect, the total PDL is not simply the sum of the contribution of each PDL element. This is due to the fact that the global PDL depends on the relative orientations of the PDL axes at each connection (The PDL axis is defined as the axis on the Poincare sphere corresponding to the polarisation state with minimum attenuation). Since the principal polarisation state of the single-mode fibres connecting the PDL

elements fluctuates with time, the relative axes of the PDL elements do also fluctuate. Hence, a statistical description of the global PDL is important, since it allows to predict the statistics of the global attenuation.

In this contribution, we report the descriptions of the mean, the standard deviation and the probability of the global

PDL of two or more concatenated PDL elements. Experimental results have been perfonned on polarising optical

fibres used as PDL elements.

PDL is defined as the maximum change in insertion loss of an optical fibre component as the input polarisation is varied over all possible states. For completeness we recall some notations and results of ref [1]. Let Tmin and T^ax denote the minimum and the maximum transmission coefficient, respectively. Then Tdepoi = (Tmax+Tmin)/2 is the transmission coefficient for depolarised light and PDLdB = 10 log (Tmax/Tmin) is the PDL expressed in dB. Since dB are not convenient "units" for fluctuating losses such as PDL, we prefer to represent PDL as a 3-dim vector f of length r=

(Tmax-Tmin)/(Tmax+Tmin) and direction corresponding to the polarisation state (as represented on the Poincare sphere) of minimum attenuation. With such notations, the transmission coefficient for a polarisation state Mj^ is given by [1] :

T=Tdepoi (l + MinT ). In [1] we showed that the PDL r,2 of two concatenated elements each having a PDL of f|

and r2 , is given by :

Introduction

Theory

(1)

163
From this equation one can derive the statistics of Fii. In particular the mean PDL M(ri2 ), assuming random relative

orientations of fj and , can be expressed with the following integral (where T) represents the cosine of the angle

between F, and F, ):

M(r,j

9 J-V fi + r,r,r|j-

(2)

Figure 2 shows theoretical plots of the mean of M(F|2 ), the mean plus and mean minus one standard deviation of Fi2 1 the maximum of F12 and the minimum of ri2 as function of F2 for a fixed F] . Note that the standard deviation

is higher for F2 close to F], and decreases when the difference F2 - Fj increases. Figure 3 shows analytical and

numerical simulation of the probability distribution of Fi2- Note that the unfavourable case of large values of ri2 is more likely to happen than the favourable case of smaller values. For more than 2 PDL elements, analytical computation becomes impractical. However, by using a simulation program

one can calculate the global PDL and its statistics. In figure 4 the simulated probability distribution of the resultant PDL of five concatenated PDL elements as function of the global PDL Fi .5 is depicted.

Experiment

The measurement principle is based on searching the maximum and the minimum transmission through the device

under test (DUT) while scanning all possible input polarisation states [2] (linear as well as circular or elliptical). The

light coming from the 1300 nm Laser is launched in a 2 m long patchcord connected to an automatic polarisation

controller, and then guided via a second patchcord to the DUT before entering a third patchcord connected with to the Power Meter. In the case of one PDL element (DUT 1) only one polarisation controller is needed. For the measurement of two concatenated PDL elements a second polarisation controller is needed (fig. 1). The DUT's used were polarising optical fibres of about 10 cm length .

1 .3 |j.m Laser

Automatic Polarisation Controller

Optical patchcord

Power Meter

Optical

patchcord

DUT 2

ax

Optical patchcord

Connector

[] DUT 1(

Connector

-S

Manual polarisation Controller Figure 1: Experimental set-up

Results

Concatenation of two PDL elements

To determine the statistics of the PDL resultant from the concatenation of two PDL elements, it is necessary to introduce a second manual polarisation controller between the two PDL elements, this polarisation controller permits

to rotate the axes of the second PDL element with respect to the first one. By rotating the manual polarisation

controller between measurements, one may obtain the average and standard deviation corresponding to on all possible orientations between the two PDL vectors. (The PDL of all optical accessories are very low and have been

neglected). Our results are summarised in table 1. the agreement with the theoretical formula (2) is excellent.

164
r,

Theoretical Mean and

Standard deviation of

Mean Sd

Experimental Mean

and Standard deviation of Fn

Mean Sd

0.102 (0.89 dB) 0.225 (1.99 dB) 0.239 (2.11 dB) 0.054 0.222 (1.96 dB) 0.053

0.184 (1.62 dB) 0.188 (1.65 dB) 0.244 (2.16 dB) 0.08 0.23 (2.03 dB) 0.076

0.744 (3.33 dB) 0.759 (8.63 dB) 0.833 (10.4 dB) 0.162 0.82 (10.04 dB) 0.150

Table 1: Theoretical and experimental results of the concatenation of two PDL elements.

Concatenation of five PDL elements

For the measurements of global PDL of 5 PDL elements, we maintained the same configuration than for the measurement of two PDL elements, except that the manual polarisation controllers were replaced by standard optical

fibre patchcords. For each measurement, the layout of the patchcords was changed in order to rotate the polarisation

and, hence to scan the relative axes of each PDL element. Our results on 65 experimental measurements are shown in table 2. For comparison we have developed a software simulation based of repeated application of formula (1). In

figure 4 the results of 65 experimental measurements show good agreement with the theoretical fit, obtained by an

iteration of 30000 simulations of fonnula (1) .

r, F, F3 F4 F5

Global PDL

Mean'" Mean'^^P Sd'" Sd^'"

0.19

(1.67 dB )

0.164

(1.43 dB)

0.178

(0.]56dB)

0.186

(0.163 dB)

0.164

(0.143 dB)

0.351

(3.18 dB)

0.350

(3.17 dB)

0.128 0.124

Table 2 : Theoretical and experimental results of global PDL resulting from the concatenation of five PDL elements,

(ih.: theoretical, Exp. : Experimental).

Figure 4 shows the experimental probability distribution of global PDL versus the global PDL of five PDL elements.

Simulation of 30 PDL elements

For large number N of devices with small PDL Fj, the global PDL derived in [2] is f = j,r+0(r"). Hence F

becomes F = M(Fj )y/N . A numerical simulation of the global PDL of up to 30 concatenated devices with 0.05

individual PDL was made. The simulated global PDL was drawn as function of the PDL number and was fitted to r = M(Fj)VN infi sure 3.

Conclusion

Statistical measurements of concatenated PDL elements have been presented and are in good agreement with the theory of referenece [1]. Regarding the difficulty of deriving an analytical expression for more than 2 PDL elements,

a simulation program has been developed. Hence, we could compare our measurements of the concatenation of up to 5 PDL elements, and support an approximate formula valid for several elements with a low PDL.

Acknowledgement:

Thanks is due to F. Prieto for his collaboration.

References:

[ 1] N. Gisin, Statistics of polarisation dependent losses, Optics communication 1 14 (1995), p 399-405

[2 ] L. Ducos, R. Clin, P. Blanchard and F. Gauthier, Polarisation dependent loss of passive optical components,

OFMC 95. 25-26 September 1995, Liege, Belgium.

165

Novel Method for Polarization-maintaining

Fiber Launch

Bernhard SchoU

Jens Rasmussen

Institute for High Frequency Technology

Aachen University of Technology

Abstract

A novel method is described to determine the principal axes of the polarization-maintaining

fiber (PM-fiber). This technique is based on the measurement of the MueUer-matrix of the

PM-fiber, but allows also a direct automatic alignment of the proper polarization.

A polarization maintaining fiber is designed to deliver a linearly polarized lightwave signal

at its output. Therefore linear polarized light has to be launched in the direction of one of

the principal axes of the fiber. Finding this direction is tedious and time consuming. Several alignment methods have been reported, one using an elasto-optic modulation method,

another using a tunable laser, both of which give an indication of misalignment of the polarization [1, 2]. In this paper, we describe a novel method which determines the direction

of the principal axes of the fiber and which includes a polarization transformation setup to

align the incident beam to the proper polarization.

The MueUer-matrix of a PM-fiber can be described as combination of a linear retarder with

the Mueller-matrix M/jet and a rotator with Mfiot

1 Introduction

2 Principle

MpM = MRot{^c)'MRet{oc,(fl) (1)

with

(\

(2)

167
and

I \ 0 0 0 \

0 cos(^c sin<^c 0

0 — sin </?c cos (/7c 0

\ 0 0 0 1 /

(3)

The angles a and </?/ correspond to the birefringence axis of the linear retarder and its phase delay, respectively, and the angle ^pc to the phase delay of the rotator. Simultaneously,

OL is equivalent to the angle of the principal axis at the input of the PM-fiber, and the

principal axis at the output is given by a -|- yjc- Knowing the Mueller-matrix MpM of the

PM-fiber both angles can be calculated easily from its components by

tan2Q; = —

m43

and

tan 2v?c m24 • "^43 - ^34 • "^42

(4)

(5)

Thus, linear polarized light with an azimuth angle of a has to be launched into the PM-fiber.

There are different ways for the determination of the Mueller-matrix of an optic device, the

setup we used adjust any desired polarization state, too.

3 Realization

The setup to analyze the Mueller-Matrix of the PM-fiber consists of a feedback controlled

polarization transformer and a standard polarization analyzer both controlled by a PC (s. Fig. 1).

The actual polarization transformation is done by three piezoelectric retarders [3]. The

piezoelectric retarders consist of a PM-fiber wound around a piezoelectric cylinder. Applying

a voltage to the cylinder lengthens the fiber and causes a phase delay between the two

principal modes of the fiber. With three piezoelectric retarders every polarization state can be adjusted if the principal axes of the three PM-fibers pieces are tilted by 45° to each other.

To control the polarization a standard 99:l-coupler follows the polarization transformer. At

the low power output port of this coupler a standard four detector photo-polarimeter is used. Therefore the light is split into four dilferent beams, which are analyzed by diflFerent polarization filters. In a calibration process the standard polarization analyzer is placed at the high power

output port. Thus, a 4x4 matrix can be found which connects the four detector signals of

the low power port to the Stokes-vector of the light at the high power port [4]. Hence, the

PC indicates the polarization at the high power output port after the calibration process.

To determine the principal axes of the unknown PM-fiber, the fiber is connected at the

output of the feedback controlled polarization transformer and the standard polarization

analyzer is placed at the output of the PM-fiber. Launching four different and known

polarization states into the PM-fiber, adjusted by the piezoelectric retarders, the Muellermatrix of the PM-fiber can be calculated by comparing the input and output polarization

168
Personal

Computer

Photodiodes ^

p>^

Polarizers @

Ay4-Plate m X /

1%- Output!

Piezol Piezo2 PiezoS ^^^^-^^^P'^'^OutpJt

Feedback controlled polarization transformer

PM-fiber

Polarization

analyzer

Figure 1: Setup for the determination of the principal axes of the PM-fiber

states of the PM-fiber. With eq. (4) the direction a of the principal axis can be calculated. In a second step the feedback controlled polarization transformer can be used to adjust the

proper polarization into the PM-fiber. This can be done easily in comparing the current

polarization state at the output of the feedback controlled polarization transformer with the

desired polarization state. For the automatic alignment of the polarization the polarization

transformer needs less than three seconds.

4 Conclusion

We have shown a possibility to determine the principal axes of a PM-fiber and to ahgn

automatically the proper polarization into the PM-fiber without changing the setup.

References

[1] S. L. A. Carrara, B. Y. Kim, H. J. Shaw, 'Elasto-optic alignment of birefringent axes in polarization-holding optical fiber', Optics Letters, Vol. 11, pp. 470-472,

1986

[2] A. Ebberg, R. Noe, 'Novel high precicion alignment technique for polarization maintaining fibres using a frequency modulated tunable laser', Electronics Letters, Vol.

26, pp. 2009-2010, 1990

169
[3] R. P. Tatam, C. N. Pannell, J. D. C. Jones, D. A. Jackson, 'Full polarization state control utilizing linearly birefringent monomode optical fiber'. Journal of Lightwave

Technology, Vol. 5, pp. 980-984, 1987

[4] B. SchoU, T. Stein, A. Neues, K. Mertens, 'In-line fiber optic polarimeter with a 99% coupler'. Optical Engineering, Vol. 34, pp. 1669-1672, 1995

170
CHARACTERIZATION OF HIGH-SPEED AND LOW-LOSS POLYMER

OPTICAL FIBERS

Yasuhiro Koike, Eisuke Nihei and Takaaki Ishigure

Faculty of Science and Technology, Keio University, Yokohama 223, Japan

Kanagawa Academy of Science and Technology, Yokohama 236, Japan

1. INTRODUCTION

The growing interests have been focused on high speed optical fiber communication in access and premises areas as well as long-distance communication. However, since the small core (5-10 \im diameter) of the single-mode fiber requires highly accurate connectors, serious

increase of the cost of the whole system is expected in such access and premises areas. The

large core (such as 300 ^im or more) of the polymer optical fiber (POF) would make it possible to adopt injection-molded plastic connectors, which dramatically decreases the total cost of the system without any complicated lens alignment.

We have proposed a high bandwidth graded-index polymer optical fiber (GI POF),

and have demonstrated several giga bit per second (Gbit/s) transmission in the 100 m GI POF

link'-^. However, the transmission distance was limited to 100 m due to the intrinsic

absorption loss. In this paper, the characterization of new GI POF with low attenuation even

at 0.5 - 1.3-fxm wavelength region is in detail described.

2. CHARACTERIZATION OF BIT RATE

In this paper, we focus on the possibility of high speed transmission by the fluorinated

amorphous polymer base GI POF. The bandwidth of the GI POF was investigated by

measuring the output pulse distortion by time domain measurement method'. As the light

source, InGaAlP laser diode with 660-nm wavelength was used. Input pulse whose full width half maximum (FWHM) is 80 psec was coupled into 100-m length POF via objective

lenses with 0.5 numerical aperture (NA) which is larger than that of NA of POF. Figure 1 shows the waveform of output pulse from the partially fluorinated polymer base GI POF'-^

compared with the conventional step-index (SI) type POF with 0.5 NA. The bandwidth of

the SI POF estimated by the result in Fig. 1 is less than 100 MHz for 100 m fiber, while in the

case of the GI POF, the bandwidth is approximately 10 GHz. A low NA SI POF has been

newly developed and 156 Mb/s transmission for 100 m link was demonstrated"*. The GI POF

will be required for higher bit rate and longer distance than those of the low NA SI POF.

Optimization of the refractive index profile of the GI POF should be the key

technology in order to minimize the modal dispersion. The possible bit rate in GI POF link

171
was theoretically calculated considering both modal and material dispersions by using WKB

method^. The material dispersion of the polymer matrix was estimated by measuring the

wavelength dependence of the refractive index of polymer^. The results are shown in Fig. 2.

It should be noted that the material dispersion of the perfluorinated (PF) polymer is much

lower than that of silica in the range of 0.4 - 1.0-^m wavelength region, which indicates that

the PF polymer base GI POF can have high bandwidth in the above wavelength region.

(A) SI POF

Input pulse

(B) GI POF

Input pulse

Ins/div.

Fig. 1 Pulse broadening through GI and SI POFs.

Fiber length: 100m

0.2 r-"—I—I—

I—I—I I

I

I—I—I—

I—I—I I I

I I I

I—

0.4 0.6 0.8 1.0 1.2 1.4 1.6 Wavelength (|xm)

Fig. 2

Material dispersion of PF polymer compared

with PMMA and silica'^

3. CHARACTERIZATION OF TRANSMISSION ATTENUATION

The attenuation of transmission of the PMMA base GI POF is shown in Fig. 3. The

minimum attenuation was about 150 dB/km at 0.65-[i,m wavelength which was almost the

same as that of the step-index type POF commercially available.

However, the attenuation of PMMA base POF was abruptly increased from about

0.6-|j,m wavelength to the infrared region due to the absorption loss of overtones of carbonhydrogen stretching vibration.

On the other hand, it is highly desirable to construct POF network system using

commercially available LD and LED which operate in the range of 0.6 -1.5 ^im of wavelength.

Deuterated or fluorinated polymer base POF will be one of the promising candidates to eliminate the serious absorption loss in such wavelengths.

172
The attenuation spectra of perdeuterated (PD) and PF polymer GI POFs measured by

conventional cut-back method are shown in Fig. 3. It is quite noteworthy that the PF

polymer base GI POF has no serious absorption peak in the range of 0.5 - 1.3 (i,m wavelength

and the attenuation even at 1.3-^m wavelength is about 50 dB/km.

5000

4500

4000

—I—'—I——

r

- PMMA base

PD polymer base

PF polymer base

0.5 0.6 0.7 0.8 0.9 1 1.1 1.2 1.3 1.4 ^. ^ Wavelength (!J.m) Fig. 3

Experimental attenuation spectra of GI POFs.

4. CHARACTERIZATION OF MODAL NOISE

It has been concerned that the modal noise degraded the bit error rate in the case of multi

mode fiber with laser diode in the fiber-optic links. However, it is noteworthy that the large core (300 - 1000 ^m diameter) of GI POF which transmits more than 30,000 modes causes no

such degradation of bit error rate even if a laser diode with high coherency is used. We

investigated the modal noise effect on the bit error rate in GI POF link as follows: a FabryPerot LD at 0.644 ^im with 1 nm spectral width was used as the light source, and the bit error

rate of 156 Mb/s system in which one fiber-to-fiber joint had been deliberately misaligned was

measured.

Figure 4 shows the results of PMMA base GI POF with a 600-^im core diameter.

Even when 200-|im misalignment was occurred, no significant degradation was observed.

On the other hand, in the case of the conventional GI glass fiber whose core diameter was 62.5

^im, it was observed that a 10-^im of displacement caused large bit error rate degradation, and

that it was impossible to obtain the accurate bit error rate curve in the case of 20-^m

displacement because of a serious fluctuation of the output power from the fiber.

It is concluded that the large core of the GI POF offers low modal noise as well as

173
low coupling loss. The large core of GI POF enables the usage of inexpensive connector

made by injection molding which tends to cause 20 to 50-^m misalignment.

o

t -7 w 1x10

1x10 r

1x10

fiber

-25 -24 -23 -22 -21 -20 -19 -18 -17-16 -15 -14 -13 -12-11 Received Power (dBm)

Fig. 4

Modal noise effect on bit error rate in

PMMA base GI POF link.

Misalignment:

GI POF

: 0 \xm •:100 ^im A:200 ^im

GI glass fiber

: 0 (im •:10 ^im MiM : 20 ^m

5. CONCLUSION

The characterization of the POF in high-speed data communication is discussed. The

perfluorinated polymer base GI POF has many advantages: one is no serious absorption peak

in the range of 0.6-1.3-^m wavelength in its attenuation spectrum, and the attenuation at 1.3

|im where the LD for silica fiber operates is about 50 dB/km. Another is the low material

dispersion compared with PMMA and silica, which allows more than 10 Gb/s transmission in

1 km link. In addition, no modal noise in the GI POF link is observed. Since the PF

polymer is very flexible and stable for many chemicals and is not flammable, outdoor use of

the PF polymer base GI POF will be possible as well as silica fiber. Therefore, we believe

that these experimental and theoretical aspects suggest that the PF polymer base GI POF will

be one of the promising candidates for "the last one mile".

REFERENCES

1. Y. Koike, T. Ishigure, E. Nihei, IEEE J. Lightwave TechnoL, 13, 1475 (1995)

2. T. Ishigure, E. Nihei, S. Yamazaki, K. Kobayashi, and Y. Koike, Electron. Lett., 31, 467

(1995)

3. T. Ishigure, E. Nihei, Y. Koike, C. E. Forbes, L. LaNieve, R. Straff, and H. A. Deckers,

IEEE Photon. Technol. Lett., 7, 403 (1995)

4. S. Yamazaki, Proc. ECOC '95, Tu. L.2.5, 337 (1995)

5. R. Olshansky and D. B. Keck, Appl. Opt., 15, 483 (1976)

6. T. Ishigure, E. Nihei, and Y. Koike, Appl. Opt. , 35, 2048 (1996)

7. J. W. Fleming,/. Am. Ceram. Soc, 59., 503 (1976)

174
A SINGLE DEVICE FOR BOTH ACCURATE

MEASUREMENTS OF ANDWq IN SINGLE-MODE FIBERS

D. PAGNOUX, J.-M. BLONDY, P. ROY, P. FACQ

Institut de Recherche en Communications Optiques et Microondes

Umt6 de Recherche Associde au CNRS

123, avenue A. Thomas - 87060 Limoges - FRANCE

The effective cutoff wavelength of the LP^j mode (K^) and the mode field radius (wq) are two of the most

important parameters for characterizing monomode optical fibers. For each measurement, many techniques have

already been proposed. On the one hand, may be measured in particular by means of the near-field shape

observation method [1], the mode interference technique [2], or the bend loss technique recommended by the ITU-T

(CCITT) [3]. On the other hand, the main methods for measuring wq are the transverse offset method [4], the knife

edge technique [5] and the far-field scan method whitch is recommended by the ITU [6]. In the recent years, we proposed a new technique based on the azimuthal analysis of the far-field to quickly and

accurately measure Xq [7]. In this paper, we show how this setup has been adapted to allow the measurement of wq,

while remaining as efficient as before for measuring X^. Only the source at the input of the fiber under test (FUT) and

the software for analysing the detected signal at the output of the device have to be switched in order to get X^ or wq.

The principle and the performances of this new dual characterization set up are presented.

First, let us remind that in the azimuthal far-field analysis technique, the part of the far-field pattern of the FUT which crosses a rotating sUt is transmitted to a detector by means of a large core multimode fiber (Fig.l).

FFT

spectrum analyser

Fig, 1 : Setup for the measurement of X^ by the azimuUial far-field analysis technique

The Fourier transform of the detected light gives a spectrum which analysis leads to the accurate determination

of : indeed, as a matter of fact, A^. and Aj, the amplitudes of tiie lines at tiie f^ and f] = f^ ± 2 fj frequencies

respectively (witii f^ = 70 Hz Uie frequency of die chopper at die input, and fs = 4 Hz die rotation frequency of the

slit) are related to the powers Pq i et P^ ^ carried by die LPq i mode and die LP^ j mode respectively, by :

175
D(A) = lOlog

1

01 = lOlog

2Ai

Ac-2A,

(1)

In the bend loss technique, Xq is the wavelength for which the total transmitted power is 0.1 dB higher than

the power transmitted by the pure LPqi mode. This criterion corresponds to a D(Kq) value equal to -16,4 dB. is

then simply determined by scanning the wavelength domain until D(K(^ reaches -16,4 dB.

By this technique, the reliability of the measurements is typically better than 3 nm for 10 successive measures. The use of the concave mirror to concentrate light on to the multimode fiber through the slit is justified by

the achromatism and the compactness of this configuration.

With regard to the measurement of wq, it has already been demonstrated that [8]

1/2 S(u) du

Wo =

An'

with

where

ft- oo u^ S(u) du

J - oo

S(uo) = Auf |<l>(u.v)|'

(2)

dv (3)

- <t)(u,v) is the complex amplitude of the far-field at the output of the FUT,

- u and V are the coordinates in the spatial frequencies space corresponding to the Fourier conjugates of x and y

respectively, which are the cartesian coordinates in the near field. S(uq) is actually the autocorrelation function of the far-field pattern, equal to the power detected through a narrow slit of Au width and set at the uq abscissa in the spatial frequencies space.

Thus, Wq may be measured with the device shown on Fig.2.

rotating

mirror

source

spatial filter^

neutral

filter

detector

fiber under test computer

Fig. 2 : Setup for the measurement of wq using the analysis of the autocorrelation function of the far-field

For practical calculation of (2), the curve S(u) is sampled. In addition uq, the abscissa of the center of gravity of S(u)

must be introduced when S(u) is not centered in the u coordinates range of acquisition. Finally we obtain :

176
n

1/2 n Z S(Ui) E UiS(ui)

Wo =

i = 0

(4) with ug

i = 0

(5) n n 47r'X; (Ui-u^)S(Ui) Z S(Ui)

i = 0 i = 0

The principles of the devices of Fig.l and Fig.2 have been combined in order to obtain a single set up

allowing both and wq measurements (Fig.3). The arrangement of the device of Fig.l essentially consists in

:

- putting a lens behind the slit, to focuse the light on to a large area detector, in the place of the concave

mirror and the large core multimode fiber,

- setting the FUT parallel to the spatial filter and reflecting the emerging light perpendicularly, towards

this filter, by means of a mirror.

The position or the motion of each component of the device are described in Table I. A large area detector is used in order to detect all the power transmitted through the slit of the spatial filter, taking into account the chromatic

aberration of the lens. Let us notice that the neutral filter and the chopper may be removed from the device and set between the source and the FUT when they are necessary (respectively for attenuating the light from the laser diode

when measuring wq, and for chopping the light from the monochromator when measuring Xq ). The performances of the dual measurement device are not degraded compared to those of each separated

measurement setup. For all the tested fibers,the measurements of X^. and wq are in very good agreement with those

given by the methods recommended by the ITU. Owing to the very abrupt fall of the curve D(X), Xq is determined

with an accuracy better than 1 nm, and a repeatability within a range of only 3 nm for tens of successive measures. The technique does not require any reference measurement and is also unsensitive to source power level and detector

sensivity drifts. It can be used for any length of fiber as well as for installed fiber links. Concerning wq, with the rotation frequency of the mirror equal to f^^ = 0.25 Hz, a curve of S(u) is stored in 200 ms and 15 successive measurements can be realized in 1 minute. The very short duration of every measurement

Fig. 3 : Dual measurement setup

177
largely reduces the effects of fluctuations of the launched power and of the sensitivity of the detector. The standard

deviation over more than 100 measures remains smaller than 0.02 |im, i.e. 0,4%. The discrepancy between the mean

value and the value given by the ITU method is typically lower than 0.03 |im for standard monomode fibers. More

detailed results will be presented during the conference.

measured parameter Wq

source laser diode at the working wavelength Xq monochromator

chopper stopped rotating at the f^, frequency

mirror rotating at fj„ fixed at the position = 45°

rotating spatial filter fixed at the position 85 = 0°

the slit is set horizontal

rotating at the f^ fiequency

lens large enough to collect all the light transmitted through the slit

neutral filter set to avoid the saturation of the detector removed

detector large area detector computer sampUng of the signal spectral analysis of the signal

and computation of relation (4) and computation of relation (1) for determining D( A.^^ = -16,4 dB

Table I : Positions and motions of the different components of the proposed device, versus the measured parameter

References

[1] Y. MURAKAMI, K. KAWANA and H. TSUCHIYA : "Cutoff wavelength measurements for single-mode

optical fibers". Applied Optics, 1979, 1 8, pp. 1 101-1 105.

[2] K. ABE, Y. LACROIX, Y. CAI and L. BONNEL :"Determination of the LPj jmode cutoff wavelength from the modal interference pattern", Proceedings of Symposium on Optical Fiber Measurements, sept 15-17 1992, BOULDER

(CO) USA, pp 221-224

[3] Y. KATSUYAMA, M. TOKUDA, N. UCfflDA and M. NAKAHARA : "New method for measuring V-value of

a single-mode optical fibre". Electronics Letters., 1976, 1 2, pp. 669-670.

[4] J. STRECKERT, "New method for measuring the spot size of single mode fibers". Optics Letters, vol 5, n°12,

1980

[5] W.G. OTTEN, F.P. KAPRON, J.C. OLSEN, "Mode field diameter single-mode fibers by knife-edge scanning in the far-field". Journal ofLightwave Technology, Vol LT 4, n° 10, 1986

[6] W.T. ANDERSON, D.L. PHILEN, "Spot size measurements for single mode fibers - a comparison of four

techniques", /owrna/ ofLightwave Technology, Vol LT 1, n° 1, 1983

[7] D. PAGNOUX, J.-M. BLONDY, P. DI BIN, P. FAUGERAS, P. FACQ, "Azimuthal far-field analysis for the measurement of the effective cutoff wavelength in single-mode fibers - effects of curvature, length, and index profile".

Journal ofLightwave Technology, Vol LT 12, n° 3, 1994

[8] B. GOULLEY, M. CLAPEAU, P. FACQ, "mode spot size measurements in non-circular optical waveguides",

£/ecf/-omc^Le«erj, Vol 22, n°18, 1986

178
Wavelength Dependence of Correction Factor of Effective Area (\eff)

and Mode Field Diameter (MFD) in Various Optical Fibers

Yoshinori Namihira

KDD R&D Laboratories

2-1-15, Ohara, Kamifukuoka, Saitama 356 Japan

Tel : +81-492-78-7801, Fax : +81-492-63-9328, E-mail: namihira@lab.kdd.co.jp

1. Introduction

Optical nonlinearities limit the maximum power transmitted through the optical fibers and they have

become a serious limitation in optically amplified ultralong transmissions spans [1]. The magnitude of

the nonlinearity in the optical fiber depends on the nonlinear coefficient of the mataial, the power in the

optical fibers and mode confinement

Recently, the measuremait of nonlinear coefficient (ni/Aeff; n2 is nonlinear refractive index, A eff is effective area) have been measured by the shelf phase modulation (SPM) method [2^], the cross phase

modulation (XPM) method in the optical fibos [5], and the self-compensated intoferometer method using

1064 nm Nd : YAG laser in the Er-doped fiber (EDF) [6] . In those methods, the nonlinear refractive index ni could not measured cfirectly. Therefore if the Aeff can be expressed approximately from the mode field diameter (MFD), it is very useful for the nonlinear measurement.

In this paper, wavelength depaidence of Ae/ and MFD for the various single-mode optical fibers is

first presented.

2. Effective Area (Aeff)

Effective area (A eff) is a parameter that is closely related to optical fiber nonlinearities that will affect

the transmission quality of the optical fiber systems, especially in long-haul optically amplified systems.

The effective aieaAeff is given by [7]

(1)

Aeff can be calculated from the measured refractive index profiles of the optical fibers [6-8] and using

the refractive index profiles as the input to computer program that models the optical transmission

properties of die fibers. Here, E (r) is the elecfromagnetic field distribution of the fundamental mode of the fiber at radius r [7]. The integration is over die entire cross-sectional area of the fiber. For example, if we make a guassian

approximation such that [8,9];

E(r) = exp (- r2/w2) (2)

where 2w is the mode field diameter (MFD), tiien eq.(l) can be integrated and gives

Kjf = 7t (3)

The Gaussian approximation is accurate for 1.31 pm optimized single mode fiber (SMF; ITU-T

G.652) and cut-off shifted fiber (CSF; ITU-T G.654) with step-index refiactive profile near the LPi i cut-off

wavelength, but for ITU-T G.652 and G.654 fibers at much longer wavelengths, and in the case of

dispersion shifted fiber (DSF; ITU-T G.653) witii non step-index refiactive profile, A^g can not be

accurately estimated from eq.(3). A more general but emprical relationship between A^ff and MFD (=2w) is [9];

179
keff = kTt (4)

where, k ( < 1 ) is the correction factor of A^^ and w. The MFD (= 2w) is defined by Petermann n

definition as shown in Recommendation ITU-T G.650 (Defmitons and test methods).

3. Wavelength Dependence ofCorrection Factor k of Aef/and MFD

In the experiment, the MFD was measured by the variable aperture test method. From the far field

pattern (FFP) of the output optical power, it was then possible to calculate the near field pattern (NFP)

using an inverse Hankel transformation. A^^is then derived from the NFP by using eq.(l).

The correction factor k in eq.(4) depends on the wavelength and on fiber parameters such as refractive

index profiles, MFD and zero-dispersion wavelength.

Fig.l, 2 and 3 show the examples of measured wavelength dependence of MFD, A^ff and correction

factor k of the 1.31 |im optimized Single Mode Fiber (SMF: ITU-T G.652) and Dispersion Shifted Fiber

(DSF: rrU-T G.653). Figs. 4 and 5 give examples of the calculated and measured results for a variety of

fiber types {SMF, Cut-off Shifted Fiber (CSF: ITU-T G.654) and DSF). From Fig.5, it was found that

the correction factor k was dependent of the refractive index profile caused by optical fiber fabrication

techniques (such as VAD and MCVD etc.,). Accordingly, the summary of the measured correction factor k of k^ff and MFD (=2w) based on the examples of Fig. 5 are shown in Table 1

4. Conclusion

The wavelength depaidence of correction factor k of P^ff and MFD was first presented. It was found

that the correcfion factor k was (fepoident of the refractive index profile caused by optical fibo" fabricafion techniques (such as VAD and MCVD etc.,). If the refractive index profile becomes stq? index profile, the

correction factor k value approachs to nearly equal to 1.0.

Acknowledgement

The author wishes to thank Mr. M. Mori and Mr. H. Ochiai of Seiko Instniment Inc, (SII), Mr. A. Hallam of York Technology/PK Ltd, for their cooperation. Also, the author would like to tiiank Dr. H. Murakami and Dr. Y. Mimura of KDD R&D Labs, for their encouragement.

References:

[1] D. Marcuse, A. R. Charaplyvy, and R. W. Tkach :"Effect of fiber nonlinearity on long distance transmission", J. Lightwave Tech., 9, 9, pp. 121-128, 1991.

[2] R. H. Stolen, W. A. Reed, K. S. Kim, and K. W. Quoi :"Measurement of optical nonlinearity of

transmission fibers," SOFM'92, NIST, Boulder, USA, pp.71-75, 1992.

[3] K. S. Kim,W. A. Reed, R. H. Stolen, and K. W. Quoi. :"Measurement of the non-linear index of

silica core and dispersion -shifted fibers," Optics Letters, vol.19, no. 14, pp.257-259, 1994.

[4] Y. Namihira et al., :"Nonlinear coefficient measurements for dispersion shifted fibres using self-phase

modulation method at 1.55 |im," Electron. Lett., vol.30, no.l4, pp.1171-1172, 1994.

[5] A. Wada, T. O. Tsun, R. Yamauchi,:"Measurement of non-lmear-index coefficients of optical fibers through tiie cross-phase modulation, using delayed-self-heterodyne technique," ECOC'92, P.42, 1992.

[6] J. E. Meier and W. E. Heinlein:" Improved self-compensated interferometer method to measure the

effective nonlineality of an Er-doped fiber", OFMC'93, Sept. 21-22, Torino, Italy, 1993, pp.125-128.

[7] G. P. Agrawal : "Nonlinear Fiber Optics", Academic Press, 1989.

[8] J. V. Wright and E. S. R. Sikora ;'T)etermination of non-linear effective areas for dispersion shifted fibers," Symposium on Optical Fiber Measurements, NIST, Boulder, USA, pp.77-80, 1992.

[9] Y. Namihira: "Relationship between nonUnear effective area and mode field diameter for dispersion

shifted fibers". Electron Lett., vol.30, no.3, pp.262-263, 1994.

180
11

10

Sc

1.1 12 13 1.4 1.5 1.6 1.7 1.8 Wavelength (jim)

Fig.l Example of measured wavelength dependence of MFD(=2w)

of ITU-T G.652 and G.653 fibers 90

80

70

r>. 60 s

"S« SO

40 -

30 -

20

SMF (G.652) A DSF (G.653)

1.1 1.2 IJ lA 1.5 1.6 1.7 1.8 Wavelength (y.m)

Fig.2 Example of measured wavelength dependence of effective area (Aeff ) of ITU-T G.652 and G.653 fibers

<

1.04

1.02

1.00

0.98

^ 0.96

0.94 -

0.92

SMF (G.652) DSF (G.653)

1.1 1.2 IJ 1.4 1.5 1.6 1.7 1.8 Wavelength ()im) FigJ Example of measured wavelength dependence of correction factor k of Aej7and MFD(=2w) of ITU-T G.652 and G.653 fibers

181
<

II

1.04

1.02

1.00

0.98

0.96

0.94 -

0.92

[Calcurated]

Cut-off Shifted Fiber CSF

(G.654)

[Step Index]

CSF(G.6S4) SMF(G.6s2)

[Non-Step Index] • DSC-DSF [G.6S3] A SGC-DSF [0.653]

IJl^m Single

Mode Fiber SMF (G.652)

Dual Shape Core DSF [G.6S3]

Segmented Core DSF [G.653]

1.1 1.2 IJ 1.4 1.5 1.6 1.7 1.8 Wavelength (jim)

Fig.4 Example of calcurated wavelength dependence of correction factor k of A*#and MFD(=2w) of ITU-T G.652, G.653 and G.654 fibers

1.04

"I

11

1.02

1.00

0.98

0.96

0.94 -

[Measured]

[Step Index] VAD-CSF (G.6S4) VAD-SMF(G.6S2) MCVD-SMF(G.6S2)

[Non-Step Index] • DSC-DSF [G.653] A SGC-DSF [G.653]

0.92 I 1 I I I 1 I 1 1 1

I 1.1 1.2 13 1.4 1.5 1.6 1.7 Wavelength (jiin)

Fig.5 Example of measured wavelength dependence of correction factor k of Affirand MFD(=2w) of ITU-T G.652, G.653 and G.654 fibers

Table 1 Summary of correction factor k of \eff and MFD (= 2w) of ITU-T

G.652, G.653 and G.654 fibers based on the examples in Fig. 5

^~''^~->«..,\V a V e I e n g t h

Fiber types ^^^^-^^

- 1.31 \lm - 1.55 Jim

IJl^ Single

Mode Fiber (SMF)

[ITU-T G.652]

VAD 0.99 - 1.00 * 0.97 - 0.98

MCVD 0.97 - 0.98

* 0.95 - 0.96

Cut-off Shifted Fiber (CSF)

[ITU-T G.654] VAD 0.98 - 1.00 *

Dispersion

Shifted Fiber (DSF)

[ITU-T G.653]

VAD

OVD

etc..

0.94 - 0.95

* 0.94 - 0.96

* ; Optimum wavelength region

182
FIBER CUTOFF WAVELENGTH MEASUREMENT BY

MODAL SPECTRUM MODULATION

P. DI BIN, C. SIMOS, D. PAGNOUX, P. FAUGERAS, P. FACQ

DSfSTITUT DE RECHERCHE EN COMMUNICATIONS OPTIQUES ET MICROONDES,

UNITE DE RECHERCHE ASSOCIEE AU CNRS

123, AVENUE ALBERT THOMAS, F-87060 LIMOGES CEDEX FRANCE

Introduction

The cutoff wavelength Xq of a fiber is defined as the frontier between the single-mode and the bimode propagation

regions. Many methods for measuring have already been presented, among which we can list : the near field shape

observation method [1], the mode interference technique [2], the far-field azimuthal filtering technique [3] and the technique recomanded by the ITU-T [4], Most of them either suffer of low speed measurement or need complex setups or information processing.

In this paper we propose a new simple method based on a spectral analysis of the Ught issued from the fiber under test when excited by alternate odd/even pattern configurations. A brief description of this method, its experimental

validation and some typical results are presented in the following.

I Principle of the proposed method

The presented method is based on the fact that the LPj j mode launched into a fiber suffers a quite infinite attenuation in the single-mode wavelength domain, whereas its attenuation is comparable to that of the LPqi mode in the bimode regime. The principle brought into operation consists in alternately exciting the LPqi mode and the LPj ^ mode at the input of the fiber under test, and comparing the power carried by each mode at the output, versus the wavelength.

According to the criterion applied in the method recommended by the ITU-T, is the wavelength for which the suppression of the LPj j mode by means of a severe curvature induces a decrease of 0. 1 dB of the transmitted power after propagation into a 2m tested fiber (the power launched into each mode at the input is assumed to be balanced).

The setup described on Fig.l provides an easy alternate excitation of the two modes, together with the spectral

analysis of the transmitted beam, leading to a very fast determination of Xq. The light from a wide band source (LED or powerful lamp) is launched into a fiber Fl whose cutoff wavelength is appreciably lower than the one of the tested

fiber : for example about 900 nm in order to characterize single-mode fibers in the 1300 nm region. The gaussian

output pattern of this fiber is coUimated by the lens LI. A rectangular glass plate with X/IO polished surfaces and

e=5mm thickness is set perpendicularly to the collimated beam, so to be crossed by exactly one half of this beam,

inducing a phase difference with the other part equal to A(p :

Acp = 1^. e (n-1) (1)

X

where n is the refractive index of the plate ( n = 1.503 for BK7 when X = 1300 nm).

The resulting beam is then focused on to the tested fiber by means of the lens L2. When A(p is equal to 2m7i,

the launched pattern is of even type (as the LPq^ mode is) and the injection efficiency is zero into the LP^ ^ mode. On

the opposite, when A(p is equal to (2m+l)7t, the launched pattern is of odd type (as the LP^^ mode is) and the injection

efficiency becomes zero in the LPqi mode. The LPqi and LP^j modes are then alternately excited versus the wavelength.

183
Intercepted half

Fig. 1 : experimental setup

The switch from one mode to the other corresponds to a change in wavelength equal to A A,

AX=

^ . (2) 2e(n-l)

In the 1300nm region, AX is approximatively equal to 0.34 nm. At the output end of the tested fiber, the transmitted

spectrum is numerized and analysed by a computer. Let us define a modulation depth M(k) of the spectrum which is the local ratio P^^^x / ^min detected power level expressed in dB.

In an ideal case, in the bimode domain of the tested fiber, the transmitted spectrum is similar to the source

spectrum, without oscillation, and M = 0 dB. On the opposite, in the single-mode region, the LPj^ mode suffers an

infinite attenuation which leads to grooves into the transmitted spectrum with a theoretically infinite depth, assuming a

sufficiently narrow spectral bandwidth of analysis. Experimentally, as shown further, the modulation depth exhibits a

finite maximum value, due to the partial coherence of the source, the spectrum analyser resolution and the residual setup

alignment imperfections. Finally, in the cutoff region, X^. is the wavelength corresponding to a 0.1 dB decrease of the modulation depth measured in the single-mode domain.

II Characterization of the setup with a monochromatic source The setup was characterized in the visible spectrum, using the monochromatic light of a 633 nm He-Ne laser. The mode pattern between the two lenses is supplied by a fiber Fl whose cutoff wavelength is now about 550 nm. The

phase change A(p is obtained by an angular tilt of the plate around its 9y axis modifying the optical path of the

intercepted half of the beam.

A sample of the same single-mode fiber is set at the focus point of the lens L2 and the output power is measured versus the value of the plate tilt. The extinction ratio of the output power measured between two extreme

excitation conditions (even or odd launched pattern) is higher than 30 dB. This very high contrast is due to the high

coherence of the laser. This result clearly demonstrates that an odd selected pattern does not excite a significant amount

of px)wer into the LPqi mode of the fiber. It then purely excites the LP| | mode.

It is also of great importance to be sure that an even pattern does not significantly excite the LP| | mode. In order to verify this assumption, the output fiber was chosen to operate in the bimode regime at ^ = 633 nm

( = 720 nm) and the far-field pattern was analysed by means of a beam analyser.

184
a) b)

Fi g 2: Experimental intensity profiles (Arbitrary units) of the far-field at the output of a bimode fiber (X^ = 720 nm)

versus the tilt of the plate (i.e. Acp) : a) LPqi pattern (left); b) LP^j pattern (right).

Fig. 2 shows the far-field intensity pattern at the output of this fiber for different tilts of the plate, i.e. for various values of A(p. It can be seen that, for suitable values of A(p, modes with a very high purity are excited. In

particular, an even launched pattern purely excites the LPqi mode and does not couple light into the LP^ | mode, as shown on Fig 2-a. Fig 2-b confirms the previous conclusion about the pure selective excitation of the LP| | mode by a

pattern of odd type. Between these two extreme situations, both LPq] mode and LPj j mode are excited. The optical

power carried by each mode was measured, in order to verify theoretical calculations of LPqj and LPj j power excitation

coefficients in the fiber F2. The LPjj injected power was measured to be about 2±0,2 dB lower than the LPqi within

the optimal conditions. This result is in very good agreement with the value of 2 dB numerically calculated. This

confirms that even in the bimode regime, the spectrum of a broadband light source will present residual modulation

versus the wavelength of at least 2 dB. As this does not concern the monomode region, there is no effect on the criterion defined above for measuring X^.

Ill Experimental results with a broadband light source The above proposed method was applied to the measurement of the cutoff wavelength of an ordinary 1300 nm

single-mode commercial fiber. This one was first charaterized by the ITU-T recommended method giving a value of

^cAeP 1257±5 nm. The setup of Fig.l is now fitted out with an LED centered around 1290 nm, close to the expected X^, value of the

fiber under test. Its spectral width is 100 nm at -40 dB (observation is limited by the dynamic range of the spectrum

analyser). The total emitted power is around -7 dBm.

Fig. 3 indicates the modulation depth M recorded at the output of the tested fiber, as a function of the wavelength. The bimode and single-mode regimes are clearly identified, described by a quite constant value of M (=5dB

and =15dB respectively). The transition between the two regions is characterized by an abrupt slope, indicating the cutoff region. Furthermore, this curve presents a great dynamic range extended up to 10 dB (against 3 dB for the ITU-T

method), thus allowing a better accuracy in determination.

In the monomode region, the modulation depth should theoretically be constant, and the corresponding part of the curve should be an horizontal straight line. The slight slope experimentally obtained (straight Une A) is attributed to the evolution of the ratio of the injection efficiencies of the two modes versus the wavelength. This slope is taken into account when applying the already defined criterion for X^. determination : X^ is the wavelength for which the parallel

straight line A' plotted 0.1 dB below A crosses the experimental curve. In these conditions, the method leads to a value of X^ equal to 1260 nm very close to the X^y^gf value (Fig. 3).

185
16 —I

1175 1200 1225 1250 1275 1300 1325

Tfavelength (nm)

Fig. 3 : Modulation depth versus wavelength for the proposed method . *** Experimental data

Fitted curve --- AandA'

Conclusion

We presented and validated experimentally a promising novel method for the measurement of the cutoff wavelength of a fiber. One of its basic advantages can be related to the controlled mode excitation conditions into the

fiber, insuring accuracy and reliability of the measurements. The method is self referenced, the measured parameter M(X)

being insensitive to source power fluctuation. The accuracy of the proposed method is now as good as the ITU-T one. Further improvements should improve these first results. The setup is small-sized, easy to utilize, has no moving parts and then permits very fast measurement (typically no more than few minutes, including the time spent for installing the fiber to test). This method is well adapted for both

fast sytematic mass production control and precise laboratory measurements.

Acknowledgements

The authors are grateful to J.F. BAYON from French CNET Lannion for providing fibers used for the setup

calibration.

REFERENCES

[1] Y. MURAKAMI, K. KAWANA and H. TSUCfflYA : "Cutoff wavelength measurements for single-mode optical

fibers", App/. Opt. , 1979. 18, pp. 1101-1105.

[2] E. BRINKMEYER and S. HECKMANN : "Cutoff wavelength determination in single-mode fibers by mode

interference". Opt. Lett., 1984, 9, pp. 28-30.

[3] J.M. BLONDY, A.M. BLANC, M. CLAPEAU and P FACQ : "Azimuthal filtering technique for effective LPll

cutoff wavelength measurement in optical fibres". Electron. Lett., 1987, 2 3, n°10, pp. 522-523.

[4] Y. KATSUYAMA. M. TOKUDA, N. UCHIDA and M. NAKAHARA : "New method for measuring V-value of a single-mode optical fibre". Electron. Lett., 1976, 1 2, pp. 669-670.

186
A SIMPLE, FAST AND ACCURATE MEASUREMENT

OF THE MODE FIELD RADIUS : THE AXIAL GAP METHOD

J.L. AUGUSTE, J.M. BLONDY, D. PAGNOUX, M. CLAPEAU, P. FACQ

Institut de Recherche en Communications Optiques et Microondes

URA CNRS N° 356

123, Avenue Albert Thomas

87060 - LIMOGES CEDEX - FRANCE

Abstract : A simple set-up is developped to measure the mode field radius (MFR) of monomode

fibers using the loss property of an axial gap.The measured MFR agrees with Petermann's definition.

At least, only two measurement points along the axis are necessary, bringing a quick and simple tool. The results are in very good agreement with those obtained with the far-field scanning method

recommended by the ITU-T.

Introduction

:

Mode field radius is one of the most important parameters of monomode fibers. Its knowledge allows the theoretical evaluation of connection and microbending losses. It is also

involved in the modelling of the chromatic dispersion. Recent experiments in the measurement of non

linear effective area [1] have increased the necessity of an accurate determination of this parameter.

Many techniques have been developped to measure the MFR. The ITU-T

recommendation uses the far-field angular scanning [2], coupled with the second definition of

Petermann. The variable aperture method [3] and the transverse offset technique [4] give alternative

ways to determine the MFR. But these techniques are time consuming specially when the field

distribution deviates from the classical Gaussian form. This drawback increases when the MFR is measured versus the wavelength for chromatic dispersion evaluation. A faster and more accurate

angular technique using a rotating mirror combined with the autocorrelation function of the far-field

gives precise, quick and reproductible measurements of the MFR [5]. But, as the other techniques, it requires the acquisition of the whole transverse far-field.

The possibility of a measurement by an axial scanning of the far field had not yet been

totally investigated. A technique has already been proposed by Saravanos etal [6], but it requires the

acquisition of the far-field along the whole axis.

In this paper, we present an axial analysis of the far-field leading to the determination of

MFR by a two points measurement. The theoretical assumptions are developped from the basis of the

loss formula for an axially offset joint between two identical fibers. A complete exploitation of the

technique shows that the MFR measured with this technique corresponds to Petermann's second

definition in both cases of Gaussian and non-Gaussian fields.

Theoretical Assumptions:

In the case of a Gaussian mode propagating in the input fiber, the expression of the

transmission factor for a longitudinal offset e is ( ignoring Fresnel losses)

:

187
T^ = —-— (1) with z=

® ^ knw^o Where n is the cladding index and Wq the MFR.

This formula is deduced from the following expression given by many authors [7], [8],

[9]: _ 1+4 z^ ^ z2.(l.2z2r

Eq(l) is obtained when noticing that : ( 1 + 4 z^) (1 + z^) = z^ + (1 + 2 zV

1 Therefore (1) is equivalent to : —— = 1 + (3)

n^ w"^

0

The inverse of the power transmission factor is a linear function of the squared axial

2

offset e as depicted in figure 2. So, two points are sufficient to determine the MFR from the slope of

the curve.

If Fresnel reflections at each endface of the fiber have to be taken into account, the new

power transmission factor is given by the following expression.

1 (n + 1)"^

16

1.

k^ n^ w^

o

(4)

Experimental set-up and results

In the experimental set-up shown in figure 1, the used light source can be a laser diode or a monochromator, or a tunable laser diode. An axial translation stage controlled by a precision stepper

motor allows the variation of the axial gap e. At e = 0 position the two fibers are aligned in V grooves

with a three axis precision translation stages. The detected power is collected by a small length of the

tested fiber and carried out to the detector.

Two different procedures can be equally used

- 2,5 m of tested fiber prealably stripped of its coating on some centimeters at a 2 meter

length are set in the V grooves. The reference power is measured. Then the fiber is cut and the axial

stage translates axially the tested fiber at different gap positions e. The mode field radius Wq is computed using eq. (4).

- the fiber is cut in two parts of 2 m for the tested one and some centimeters for the

connection to the detector. The two fibers are aligned to obtain the maximum power and then

translated to different gap position e. The MFR is computed using eq (3) because the Fresnel

reflections are present for both reference and power measurements.

188
Chopper

Light

source

2m tested fiber

Detector

Translation

stages

Lock-in

Amplifier

Figure 1 : Experimental set-up

2 . As shown in figure 2, the curve 1/^ versus e is a straight line. In the first micrometers

range, the measurement is perturbated by a Fabry-Perot cavity effect. That is why it is better to make

the measurement with a wide gap between the two fibers (from 2 to 5 millimeters in our experiment).

Fi gure 2 : Experimental curves of the measured inverse power transfer coefficient 1/T

2 versus the squared axial gap e

189
The average value of MFR measured with the axial technique for a standard SMF 28 fiber

is 4.54 (j.m very close to the value obtained with the ITU-T recommendation ( 4.55 )J.m ).The

discrepancy between ten successive measurements is less than 1%.

Conclusion:

The axial gap method gives a powerful tool for the measurement of the mode field radius.

Two measurement points along the axis are sufficient. This leads to one of the fastest methods to obtain Wq. Our first experiments show a good agreement of the measured values with the ITU-T

recommended far-field angular scanning method (within some 10 |im). The influence of different

parameters such as precision on the gap e, power drifts of the source, small angular tilt of the fiber,

uncertainties about the reference power, and linearity of the detector are being investigated.

References :

[1] Y. NAMIHARA, "Relationship between nonlinear effective area and mode field

diameter for dispersion shifted fibers". Electronics Letters, Vol 30, n°3, pp 262-264, 1994.

[2] Recommendation of the CCITT n° G 652

[3] F. ALARD, L.B. JEUNHOMME, P. SANSONETTI, "Fundamental mode spot size

measurement in single-mode optical fibres". Electronics Letters, vol 17., n°25, pp 958-960, 1981.

[4] J. STRECKERT, "New method for measuring the spot size of single mode fibers".

Optics Letters, vol 5, n°12, pp 505-506, 1980.

[5] J.M. BLONDY, D. PAGNOUX, J. RIOUBLANC, P. ROY, P. FACQ, M.

CLAPEAU, "Fast and accurate measurements of the mode field radius using the autocorrelation

function of the far-field". Proceedings of OFMC95, 25-26 septembre 1995, LIEGE (Belgique).

[6] C. SARAVANOS, R.S. LOWE, " The measurement of non-gaussian mode fields by

the far field axial scanning technique" Journal of Ligthwave Technology, vol. LT4, n° 10, pp 1563-

1566, 1986.

[7] D. MARCUSE, " Loss analysis of single mode fiber splices". Bell System Technical

Journal, vol.56, n°5, pp 703-718, 1977.

[8] L. B. JEUNHOMME " Single mode fiber optics" 2nd edition Marcel Dekker inc.,

New-York and Basel, 1990.

[9] C. M. MILLER, S.C. METTLER, I. A. WHITE, "Optical fibers splices and

connectors". Marcel Dekker inc.. New-York and Basel, 1986.

190
Calibrated Test Fiber for

Optical Time-Domain Reflectometers

Duwayne Anderson, Tektronix, inc.

Calibration of optical time-domain reflectometers (OTDRs) has traditionally involved calibration

of the instrument's vertical and horizontal scales. For early OTDRs this level of instrument

calibration was sufficient because these early machines provided only a waveform, with some

cursors and zoom functions to assist the operator in making the actual measurements.

Modem OTDRs not only display the standard waveform, they also generate event tables that list the locations, losses, and reflectivities of components and anomalies located on the optical fiber.

Mathematical algorithms, which interpret the OTDR's waveform and make measurements on it, generate these tables. Waveform characteristics such as noise, recovery from reflections, and

digital filtering affect the accuracy of these algorithms. Tests show the system-level repeatability

of many OTDRs is an order of magnitude less accurate than the hardware calibration of their

subsystems [1]. Thus, it is no longer sufficient simply to calibrate the OTDR's hardware

elements. We must also calibrate (or at least characterize) the OTDR's system performance,

including absolute measurement errors arising from deficiencies in the interpretive algorithms.

A delay generator can be used, in conjunction with an optical-to-electrical converter and an

optical transmitter, to calibrate the OTDR's time base and horizontal scale [2]. An optical

transmitter and a variable attenuator can be used to calibrate the OTDR's vertical scale. Both of

these subsystem-level calibration tasks involve non-fiber equipment that is easily calibrated to traceable standards. Calibration of the OTDR's system-level performance, however, requires a

fully calibrated optical fiber with real components (events) such as connectors and splices,

located along its length.

Three quantities describe each event on an optical fiber: distance from the OTDR, optical loss,

and reflectivity. Ideally, the distance, loss, and reflectivity of each event on the test fiber should

be calibrated and fully adjustable. We have developed such an optical-fiber test fixture that has

three calibrated events. The first two events have calibrated distances, but not calibrated

reflections or losses. The third event has calibrated distance, loss, and reflectivity, and the loss

and reflectivity are both fully adjustable (see table 1).

Figure 1 illustrates the test fixture, which consists of the following components:

1. Four rolls of fiber. The lengths of the fibers are measured by the standard time-of-flight

method, which gives a precise length for the fiber at a given (assumed) group index. The

fiber loss is measured at 1310 and 1550 nm using the standard cut-back method [2]. 2. Two calibrated optical attenuators with low insertion loss. 3 . An optical coupler with measured split ratio at 1310 and 1550 nm.

4. A pigtailed mirror with measured reflectivity.

5. A device for bending the optical fiber and introducing loss.

6. A calibrated optical power meter.

191
= splice = connector

Figure 1 . Configuration of the optical test fiber. The length of each section of fiber is measured with the time-offlight method, and the insertion loss determined by the cut-back method.

Table 1 Approximate location, type, and calibration status of events

Event Tvpe of event Event location Event loss Event reflectivitv

1 Reflective NA NA Uncalibrated

2 Low reflectance Calibrated Absolute loss

uncalibrated

Change in loss

calibrated

Uncalibrated

3 Low reflectance Calibrated Uncalibrated Uncalibrated

4 Adjustable:

reflective or nonreflective

Calibrated Calibrated and

adjustable

Calibrated and

adjustable

5 Reflective Calibrated NA Uncontrolled

The first event on the test fiber is the reflection at the OTDR's front-panel connector. The

second event is at distance Lj . The event's insertion loss is controlled by the variable attenuator,

but is not calibrated because the losses of the connectors are unknown. However, the change in

192
loss (as a function of the attenuator loss) is calibrated. The attenuator is useful for raising or lowering the remaining events throughout the OTDR's dynamic range. This is important

because the OTDR's performance on subsequent events can be altered depending on how close

the events are to the noise floor. Together, the first and second attenuators may be used to move

the reflection of the mirror through the OTDR's full vertical scale range. This is useful for

characterizing and calibrating the OTDR's system linearity.

The second event is the coupler (the second attenuator is physically located less than 0.5 meter

from the coupler, and is indistinguishable from it on the OTDR trace). The loss and reflectivity

of this event are uncontrolled, but the distance is calibrated. The insertion losses of all the fiber

spools are measured before assembly. Consequently, their link losses may be measured as the

fiber is moved up and down through the OTDR's dynamic range by adjusting the loss of the first attenuator. This is one of the calibration procedures proposed by the lEC. The technique uses a

reference fiber of known loss, moving it through the OTDR's vertical scale, and verifying that

the fiber's link loss as measured by the OTDR remains constant [2].

The distance to the mirror (through one leg of the coupler) is exactly the same as the distance to the fiber bend (through the other leg of the coupler)'. The reflection from the mirror is overlaid on the OTDR's display with the non-reflective loss of the fiber bend. By adjusting the loss of the

fiber bend, and the attenuator in the coupler leg with the mirror, the composite loss and

reflectivity of the event, as seen on the OTDR's display, is fully adjustable and calibrated.

The true reflectivity of the mirror (as measured by the OTDR) is a fiinction of the mirror's

absolute reflectivity, the losses of fusion splices, the coupler split ratio, and the adjustable

attenuator. The attenuator is calibrated, and the mirror's reflectivity characterized, before

assembly. The splice losses are estimated from the fusion welder, and the coupler split ratio

measured before assembly. Total uncertainty for the reflection is relatively high compared with

tolerances for insertion loss, but well within tolerances specified by Bellcore [3].

The bend loss is calibrated from readings at the optical power meter, while the OTDR is in its laser-pulse mode. In the laser-pulse mode the laser fires at a constant rate, high above the power

meter's bandwidth. With the OTDR in the laser-pulse mode, the optical power emanating from

the far end of the test fiber is measured when the fiber is not bent (see figure 1). Next, the fiber

is bent into a loop sufficient to attenuate the power meter's reading by a predetermined amount

(0.25 dB, for example). The true calibrated loss that should be reported by the OTDR is:

[1]

J

Precise assembly and calibration techniques are required to ensure the equality of these two distances.

193
In equation [1] Losspy^ is tlie loss seen by the optical power meter (the difference in the power

meter readings before and after the loop loss is introduced, in dB). The variable Q\S3. correction

factor accounting for backscatter emanating from the coupler's second leg (the one with the

mirror). If the event's reflectivity (as measured by the OTDR) is not too large, then ignoring Q

results in an error of less than about 0.001 dB (assuming the mirror's absolute reflectivity is greater than about 85%).

Distance

Figure 2. OTDR waveform of the test fiber. We have found this test fixture to be an important tool for assessing the absolute accuracy of

system-level measurements made with OTDRs. With this fixture, the OTDR's absolute accuracy

for loss, distance, and reflectivity measurements can be tested and verified over the ftall width of

the OTDR's vertical dynamic range. By attaching fibers of calibrated length to the fixture, the

OTDR's performance can also be investigated over the instrument's distance-measurement,

range.

References:

[1] Duwayne Anderson and David Judge, A Pattern Matching Algorithm for Remote Systems

that Measures the Distance and Loss of Fusion Splices with High Precision, National FiberOptics Conference, 1994.

[2] lEC TC 86 / WG4 / SWG2 Calibration OfOptical Time-Domain Reflectometers.

[3] Bellcore GR-196-CORE Generic Requirements for Optical Time Domain Reflectometer

(OTDR) Type Equipment.

194
A Method for the Localisation and Quantization of Faults in Passive

Tree-Structured Optical Networks using the OTDR technique

L. Wuilmart, V. Moeyaert, D. Daniaux, P. Megret, M. Blondel

Service d'Electromagnetisme et de Telecommunications

Faculte Polytechnique de Mons

Boulevard Dolez 3 1 B-7000 MONS (BELGIUM)

ABSTRACT

A method for the detection, localisation and quantization of faults in tree-structured passive optical networks is presented. This method is based on the analysis of OTDR (Optical Time Domain Reflectometry) traces. Several kinds of faults have been generated in a 1X8 passive optical network to illustrate the effectiveness of this technique.

INTRODUCTION

The use of OTDR has proven to be very efficient for the maintenance of optical links. The access to only one

extremity of the link is necessary; the OTDR signal wavelength and the data signal wavelength can be different so

that the monitoring of the link involves no interruptions of the communications.

If a fauh occurs in a point-to-point optical link, a simple comparison and analysis of OTDR traces leads to an

accurate determination of its position, its kind and parameters such as its loss and return loss. The problem is more complex in the case of a tree-structured network, since the trace acquired at the top of the network results in the combination of the reflected and backscattered light signals coming from the different optical

paths. If a fauh occurs in a branch of a tree-structured network, it can be easily detected comparing the trace acquired at the head of the failed network with a reference trace. However the affected branch and the fault's parameters (kind, loss, return loss) are more difficult to determine.

The method explained in this paper solves this problem. It is illustrated with the analysis of a 1X8 passive optical

network. Several kinds of faults have been generated and the results obtained with a software implementing this method are presented.

METHOD

Fault detection and localisation A fault in a point-to-multipoint optical link can be easily detected comparing an OTDR trace of the failed-supposed

network with a faultless reference trace : a significative difference between these two traces is a failure sign.

Moreover the failure is situated at the position where the trace difference becomes higher than a given threshold. The problem lies in the determination of the failed branch on the basis of the analysis of an OTDR trace acquired at the network's head, since such a trace results in the combinations of all the backscattered and reflected signals.

Nevertheless as we can identify each branch on the trace with an end-reflection peak, we can determine the failed branch which is the one whose height's variation of the end-reflection peak is the more important (see examples in the experimental study). This assumes that significant and equivalent reflections are generated at the OTDR signal

wavelength at the end of each branch and that the lengths of all theses branches are different.

Fault Quantization

The end-reflection peak's variation gives information about the kind of network failure: a left shift of a peak

indicates a cut, a small height's variation can be caused by a small bend of the optical link, ... Other examples are given in the experimental study.

When a fault occurs in a point-to-point optical link, its loss can be directly measured on an OTDR trace. But this is not the case anymore for a tree-structured network, a simple analysis of an OTDR trace acquired at the network's

head doesn't provide this parameter.

Actually to have an accurate fault loss estimation on the basis of a complex trace analysis, it is necessary to take

into account the exact contribution to the complex trace of the reflected signals coming from each branch. One often thinks that this involves a previous knowledge of network's parameters [1][2] such as the splitter losses, the fibers attenuation, ... However we show below that the fault loss can be obtained using only OTDR traces informations.

195
Two traces informations are power levels (dB) measured by the OTDR at the top of the end-reflection peak

associated to the fault branch, first on the reference trace (Pref) and second on the fault trace (Pfault).

peak/branch

I i peak branch

Pref

Pfault

n' peak/branch

fault distances

Figure 1 : fault trace (fault in the i^branch)

Considering a fault in the i'*' branch, these power levels can be expressed as the sum of the reflected signals

contributing to the trace at the i"^ peak position

Pref = 51od XlO^^*"^

Pfault = 51og|l0P«-Lf+ JiqPO^)

V k=i+l

(1)

(2)

where P(k) is the power level (dB) at the i* peak's position of the reflected signal coming from the k* branch, Lf is the fault loss. From (1) and (2) we can compute the fault loss (Lf) :

Lf = -51og

Pfault

10 -10

Pref

5 + 1

V

(3) P(i)

10 5 We see that the loss depends on P(i). Actually P(i) represents the power level that would be measured at the top of

the fault branch's end-reflection peak if only the reflected signal from this branch contributes to the trace at the peak's position. In order to perform a fault loss determination, we must know this power level for each branch. As we base on the fact that the part of a trace from the end of the penultimate branch to the end of the longest one is only composed of the reflected signal coming from the longest branch, it is possible to measure these power levels

at the network's installation. The branches must be connected to the splitter in increasing length order and a trace must be acquired after each connection. So that, for each branch, we have a trace where it is the longest one, and

therefore we can measure the wanted power level associated to each branch's end-reflection peak.

Advantages

- A main advantage of this method is its low cost compared with other reflectometric methods using active [3] or WDM [4] components inside the splitters.

- This method can apply to multistaged passive tree-structured optical networks.

- In order to estimate the fault loss, it is not necessary to previously define the network's parameters (splitter losses, fibers attenuations, ...). - The presence of significative end-reflections at the OTDR wavelength increases the fault detection sensibility: in the case of a small loss, a more important modification of the trace will be observed with than without reflection peaks.

Drawbacks

- Since the trace is acquired at the network's head, the OTDR signal suffers the splitters losses. Therefore the OTDR

dynamic -the maximum measurable loss-, becomes a limiting factor. The OTDR dynamic can be inproved in increasing the pulse duration but at the expense of the spatial resolution. However having a very accurate spatial resolution is not always needed as we suppose that faults generally appeared at well-known located connectors or

splices or are created by external factors such as public works which can be accurately located.

- The branches must have different lengths in order to distinguish their end-reflection peak.

196
EXPERIMENTAL STUDY

Experimental Set Up

The effectiveness of this fault analysis method was demonstrated by acquiring and analyzing a 1X8 splitter's OTDR

trace of a connected network build in our laboratory (see figure 2). The input fiber was about 2.50 km long and,

behind the splitter, the length of the fibers was between 2.26 km and 6.10 km. All fibers were terminated on a 4%

Fresnel reflection (right angle air/fiber interface).

2.260 km

2.530 km

2.550 km

2.695 km

2.865 km

3.095 km

3.670 km

3.990 km

6.100 km 8

Figure 2 : Experimental set up. An initial test was performed at 1550 nm using a 100 ns optical pulse. Figure 3 shows the 'installation' (reference)

trace of our network on which one can easily see the large insertion loss characteristic of the 1X8 splitter and endreflection peaks associated to each branch.

Figure 3 ; 'Installation' (reference) trace of the network - Power [dB] vs distances [m].

Faults analysis After this initial testing, some impairments were done to the network. Four are illustrated below:

1 . a bend in the 3"* branch at 1 70m from the splitter (about 7dB loss).

2. a large bend in the 5* branch at 560m from the splitter

3. a break in the 3'^'' branch at 170m from the splitter

4. a break at the end of the 5* branch

For each impairment a new OTDR trace has been acquired and let to be analysed by our software in order to assess

its quality in fault discrimination. Our software can display the reference trace, the fault trace and the difference between these traces (in the examples below, only the fault trace and the difference are displayed). A map of the network is drawn and a cross indicates the position of the fault. The fault loss, the fault position and a comment

concerning the possible type of fauh are also given.

197
! Il

r Reteience T.r

;oo 8000 80g ItOO 2400 32D0

rRefetence Trace PIotdrIPrecision f m) in bran ch Loss (dB)

Commeni ILoee without Relledmn (band, bod splice. )

Print

I |r"duk jj

Precision f m) in brnr>ch Distance Irom |567

[ [iDj js|

Comrr»ent: |possibility ol large bend Figure 4: bend in the 3' branch at 170m from the spHtter Figure 5: large bend in the 5 branch at 560m from the

splitter.

1-1*1x1

BOO uoo z^ocii' "J2U0 lootf isoo ~^60o \m ^0 ma

r Belerence Trace '

1600 im 3^00 ^OOO 4900 UOO 6400 7200

r Relerence Trace Prr.i.l+ li^- o

lOTORlDialance tror on fm) in branch Loss (dB)

Possibility ofbiBok

Print [r Qo^

Precision f m): in branch- loss [dB) Distance from [3066 1 [To fl I [m \ splitter mm '

— ' ' 1 ' '

I I Comment: iBreak or Return Lose modilication at the end ol the branch Figure 6: break in the 3' branch at 170m from the splitter Figure 8 : break at the end of the s"' branch

CONLUSIONS

The method presented showed that the optical time domain reflectometry can be used to facilitate the maintenance

of multistaged tree-structured passive optical networks at a low cost. This method has been implemented in a user-friendly software, and the results obtained proved its effectiveness. We also showed that it is possible to estimate the loss of a fault only on the basis of OTDR traces acquired at the network's head despite the fact that such traces resuh in the combination of the reflected and backscattered signals

coming from the different optical paths.

AKNQWLEDGMENTS

This work was funded by BELGACOM in the frame of RD/TR/UN/1 contract. REFERENCES

[1] F. Kapron, J.D. Berardinelli, "OTDR Measurements Through Optical Splitters", proc. Optical Fiber

Measurements Symposium, Boulder (USA) 1992.

[2] Caviglia et al., "OTDR analysis of tree-structured optical networks. Fiber test methods and field testing", proc. EFOC&N, 1994.

[3] Y-K. Chen et al., "Fault-Locating and Supervisory Technique for multistaged Branched Optical Networks",

IEEE Photonics Technology Letters, Vol.6, n°7, July 1994.

[4] F. Yamamoto et al., "In-service Remote Access and Measurement Methods for Passive Double Star Networks",

proc. 5th Conference on Optical/Hybrid Access Networks, Montreal (Canada) September 1993.

198
A COARSE RANGING MEASUREMENT METHOD FOR ACCESS SYSTEMS

USING PASSIVE OPTICAL NETWORKS

A. FELLEGARA

ITALTEL -A Stet & Siemens Company -R&D Transmission Business Unit

20019 Castelletto di Settimo Milanese (Milano) - ITALYAbstract: A coarse ranging measurement method based on the use of an

asymmetrical bidirectional system architecture is presented. The impact of the

crosstaLk on. system performances is calculated and discussed.

In these years the deplo)anent of fiber-optic cables is penetrating further to the local

loop plant to bring broad band and/or narrow beind directely to the customers. One

candidate for this network is the Passive Optical Network (PON) which incorporates

passive optical splitters, forming a point to multipoint network. In other words an

Optical Line Termination (OLT) at the central office is connected by optical fibers to a

plurahty of substations called Optical Network Units (ONUs). The OLT centrally

controls the data trafic to and fix)m the various ONUs and, in particular, the

information flow from the OLT to each ONU is called the "downstream transmission",

whereas the information that flows from each ONUs to the OLT is called the " upstream

transmission". Due to the pecuHar character of this data trafic, it is essential to have a

ranging procedure (that can be coarse or fine depending on the accuracy) which accounts

for the distances between the main station and the substations Edlowing the transmitted

data to arrive at the specified time slots for the substation with respect to the main

station timing. This paper aims to investigate a coarse ranging method based on the

use of low bite-rate optical signal traveling in downstream and upstream fibers always

counterpropagating with respect to the data trafic. This method can be made clearer by

the non exhaustive example reported in fig 1: the OLT activates the ranging transmitter

which sends a low bit-rate ranging signal counterpropagating with upstream data. The

'to-be -adjusted' ONU re-sends to the OLT the ranging, counterpropagating with the

199
downstream data. The measurement of the occurred round trip delay gives the

desired ranging value. The coupling ratio of the optical splitters aUows the transmission

of the CW and BURST signals without any significant degradation of the power

budget whereas the ranging losses are well compensated by the smaller value of the

ranging sensitivity. A feasibihty analysis of the system according to the proposed

ranging measurement can be performed assuming the power budget given by [1]

-i-=A.(k-y)'

P.-<uu (1)

P. . - X2 P* r-rang

Where both couplers have a power sphtting ratio of k:(l-k), y is the coupler insertion

loss, A is the fiber loss between the couplers, P^ and Pr-rang/data

^

^® optical

transmitted and received powers for the remging and data, respectively. The received

powers (for A=Amax) coupling ratio are shown in fig.2. As an example, by

using 10/90 couplers, the received powers are Pf-rang ~* ^^-^ dBni (10 %) and

Pr-data~*2'7.5 dBm (90 %), respectively. To evaluate the bit-rate let us impose that

Pr-rang/data satisfy also the following condition

Q

P

Pr-rang/data - ~( V*^ 'ttierm > + 2qpBPfefl (2)

where Q=6 , Ppefl is the backreflected power that reaches the photoreceiver and both

sides of eq.2 are equal when Pf-rang/data receiver sensitivities, i.e the

minima detectable average optical powers for a given BER. A simple calculation shows

that 622 Mbit/s (data) and 2 Mbit/s (ranging) are solutions consistent with eqns. (1-2)

at Prefl ~0 since the correspondent baseline sensitivities are well below the received

powers [1]. However, bidirectional direct-detection systems are sensitive to the

crosstalk fi'om Rayleigh backscattering and discrete reflections that, in PON systems,

are mainly caused by the laser diode reflections at the ONU side [2]. The multipath

interference and coherent Rayleigh noise [3], can be significantly reduced by using

multimode Fabry-Perot semiconductor lasers, whereas a ranging bit-rate much lower

than the data trafic bit-rate allows to filter out the crosstalk before the decision point on

both receivers. On the other hand, the impact of the crosstalk-induced shot noise on

200
the system performances is evaluated with the help of eq. 2, by calculating the power

penalty with respect to the baseline sensitivity. Fig, 3 reports the power penalty vs the

crosstalk-receiver sensitivity ( CR = Prefl^s-rang/data)' showing a similar behaviour

for both data and ranging curves. From this figure it can be concluded that for a wide

range of reflection levels the increase of the sensitivities is well below the received power

values corresponding to the 10/90 couplers. To understand these results, it is worth

while to note how different is the role of the crosstalk played in the conventional

bidirectional systems: in our case, in fact, the strong asymmetry in the operating bitrates allows to filter out the signgj-crosstalk before the decision point on both receivers.

Thus, the crosstalk degradation is due to the crosstalk-induced shot noise which has a minor impact on the system performemces.

In summary, a digital coarse ranging measurent method for PON systems is presented

and the system penalties induced by the crosstalk are also calculated and discussed.

REFERENCES

[1] A. Fellegara - Electron. Letters, vol n.l, p.51 1996

[2] I. Van de Vorde, C. Sierens - ECOC '95-Brussels, paper Tu.P.19 p.449 , 1995

[3]R. StaubU, P.Gysel - J. Lightwave Tech., vol 9, p.375 , 1991,

TX-RANG

DOWNSTREAM

Fig.l - Architecture of the optical link according to the proposed ranging procedure.

201
Coupling ratio (1-k)*100 [%) 85 80 75 70 65

15 20 25 30 35

Coi^iling ratio k* 100 [%]

Fig. 2 - Received Power vs Coupling ratio. Dotted /i/i€s:theoretical calculations x, + =

experimental data measuring power couplers with various splitting ratios. Data:

Amax= 26 dB, Pt= 0 dBm, y=0 04.

-20 15 -10 -5 0 5 10

Crosstalk- Receiver sensitivities (dB)

Fig.3 - Power Penalty vs Crosstalk power-receiver sensitivity. Data: p=0.73

Solid lines: 622Mbit/s, Pg= - 31.9 dBm, (data baseline sensitivity) , from [1]

Dotted lines: 2 Mbit/s, P§= - 56.9 dBm, (ranging baseline sensitivity), from [1]

202
spatial-resolution improvement in coherent optical frequency domain

reflectometry for long optical fibers by reducing frequency-sweep nonlinearity

K. Tsuji, K. Shimizu, T. Horiguchi, and Y. Koyamada

NTT Access Network Systems Laboratories

Tokai-mura, Ibaraki-ken, 319-1 1, Japan

Abstract

We report long-range coherent optical frequency domain reflectometry (C-OFDR) with an

auxiliary feedback-control circuit to suppress optical-frequency-sweep nonlinearity. We achieved

spatial resolutions of 1 m and 50 m at the far ends of 1 km and 30 km optical fibers, respectively,

without sacrificing the measurement time. These are the best results reported to date for C-OFDR.

1. Introduction

It is important to be able to measure loss and locate faults in optical fibers with high spatial

resolution. For this purpose, coherent optical frequency domain reflectometry (C-OFDR) is very

attractive because of its high spatial resolution and high sensitivity [ll-[6]. With this technique,

the frequency of a lightwave is swept linearly with respect to time, and the frequency-swept

lightwave is divided into probe and reference Hghtwaves. The probe lightwave is launched into an optical fiber under test, and the reflected-probe lightwave from the test fiber is mixed with the

reference lightwave at an optical receiver. The resultant beat signal is analyzed in the frequency

domain. The beat frequency is in proportion to the optical-path difference AL between the

reflected and reference lightwaves, as follows,

/. = r^/v, (1),

where yis the frequency sweep rate and is the group velocity of light. Therefore, the reflection

position can be identified uniquely from the beat frequency ^ as long as the frequency-sweep

rate yis kept constant. The spatial resolution is improved by narrowing the frequency-resolution

bandwidth, and the achievable spatial resolution is theoretically given by using the total

frequency-sweep span AF, as follows,

Az^,,= v^/2AF (2).

Therefore, a wide frequency-sweep span is required for high spatial resolution.

In previous reports, there have been a few experiments which employed C-OFDR for the

measurement of long optical fibers [2] -[6]. In these measurements, frequency-sweep nonlinearity

severely limits the spatial resolution. This is because it causes a fluctuation in the beat frequency

through the change in the frequency-sweep rate 7 according to Eq. (1), which results in a broadening of the beat spectrum and degrades the spatial resolution. Since the beat-frequency

fluctuation is more obvious when the optical-path difference AL is large, it becomes more

difficult to achieve high spatial resolution as the length of the test fiber increases. In a previous

study [6], we reduced the spectral broadening which results from the nonlinear-frequency sweep

by using time-gated spectrum analysis, and improved the spatial resolution. With this technique,

however, since a limited part of the frequency-sweep span and time are used, the achievable

spatial resolution degrades according to Eq. (2) and the measurement time is considerably sacrificed

depending on the degree of the frequency-sweep nonlinearity. In this paper, we report a C-OFDR in which the frequency-sweep nonlinearity itself is suppressed by employing an auxiliary feedback-control circuit. We successfully improved the

spatial resolution of C-OFDR without the substantial increase in measurement time reported in

[6].

203
2. Experimental setup

Figure 1 shows the experimental setup, which consisted of the C-OFDR apparatus and an

auxiliary feedback-control circuit to suppress the frequency-sweep nonlinearity. The lightwave

source was an Er-Yb-glass laser operating at 1 .55 |j.m with a linewidth of approximately 3 kHz,

The lightwave was sinusoidally modulated with an external electro-optical (EO) phase modulator

incorporating a Ti-diffused-LiNbOj waveguide. The EO phase modulator was driven by a synthesized-frequency sweeper. The lightwave frequencies of the modulation sidebands were

swept by changing the modulation frequency linearly with respect to time. The frequency sweep

rate 7 was approximately 0.1 THz/s, where the frequency-sweep span was approximately 2 GHz

and the sweep time was approximately 20 ms. Accordingly, the relation between the fiber length

and the beat frequency was approximately 1 m/kHz. The frequency-modulated lightwave was

divided into probe and reference lightwaves by using a directional coupler (DCl). The probe

lightwave launched into an acousto-optical (AO) frequency shifter, which separated the beat

spectra caused by the upper-modulation sidebands from those caused by the lower sidebands.

The frequency-shifted-probe lightwave passed through an erbium-doped-fiber amplifier (EDFA),

and was launched into an optical fiber under test by using a directional coupler (DC2), where the

total optical power launched into the test fiber was 5.0 dBm. The reflected lightwave output from

the test fiber was extracted through DC2, and was detected with an optical balanced-heterodyne

receiver (ORl) by using the reference lightwave as a local oscillator (LO). The total LO power

was -5.4 dBm. The photoelectric current from the optical receiver was analyzed with a spectrum

analyzer. The signal-to-noise ratio was improved by averaging the beat spectrum 100 times.

Time-gated spectrum analysis was used with a time-gate width of 5 ms. By contrast, in the feedback-control circuit, the non-frequency-shifted lightwave from the

AO was introduced directly into an optical balanced-heterodyne receiver (OR2). The frequencyshifted lightwave from the DC2 passed an isolator and an optical-fiber-delay line with a length of

approximately 500 m, and was launched into OR2, where the beat signal between the two

lightwaves was obtained from the photoelectric current. This beat signal was mixed with an RF

signal and passed through a band-pass filter (BPF) to provide a reference signal. The frequency

fluctuation of the reference signal, which was caused by the nonlinear-frequency sweep, was

obtained by using a frequency discriminator (FD). The FD output, which passed through a low-pass filter (LPF) and an amplifier (AMP), was used for the frequency-sweep control. The

control signal was fed into the FM-control terminal of the frequency sweeper.

3. Experimental results

Figure 2 shows the spectrum of the 0R2 output. The center peak is the beat signal caused

by the optical carrier and the other peaks are due to the modulation sidebands. We used the beat

signal at a frequency of 80.29 MHz as the reference signal. The frequency change in the reference

signal is obtained from the FD-output signal as shown in Fig. 3, where the signal-to-noise ratio

was improved by averaging the signal 100 times. The solid and the dotted lines in Fig. 3

represent the frequency change with and without the frequency-sweep control, respectively.

Consequently, we confirmed that the beat-frequency fluctuation was greatly reduced by using our modulation-frequency-sweep-control circuit.

Figure 4 shows OFDR waveforms for the vicinity of the far end of a 1 -km-long optical

fiber. The horizontal and the vertical axes correspond to fiber length and reflected-lightwave

power, respectively. The resolution bandwidth of the spectrum analyzer was 300 Hz. The waveforms

shown in Fig. 4(a) and (b) are beat spectra without and with frequency-sweep control, respectively.

In Fig. 4(a), the peak due to the reflection at the far end of the test fiber broadened as a result of

the frequency-sweep nonlinearity. Since the spectral width (3 dB below the peak) was approximately

4 kHz, the spatial resolution (3 dB) without the frequency-sweep control was approximately 4 m.

204
In Fig. 4(b), where the frequency sweep was controlled, there was an improvement in the spatial

resolution for the far-end reflection. With the control, the spectral width (3 dB) was reduced to approximately 1 kHz, which gave a spatial resolution (3 dB) of 1 m for the far-end reflection. It was, however, 5 times as large as the theoretical value of 20 cm given by Eq. (2). This was due

to the residual nonlinearity in the frequency sweep.

Furthermore, we measured a 30 km optical fiber. We have already shown that C-OFDR is possible, even for such a long optical fiber, by using the beat spectrum resulting from the mixing

of phase-decorrelated lightwaves [6]. Figure 5 shows OFDR waveforms for the far-end of the

test fiber. Here, (a) and (b) are beat spectra without and with frequency-sweep control, respectively.

We set the resolution bandwidth of the spectrum analyzer to 10 kHz which was twice the

linewidth of the lightwave, because the phase-correlated-beat signal for the far-end reflection

disappeared and the beat-spectrum width broadened. By reducing the frequency-sweep nonlinearity,

the spatial resolution (3 dB) for the far-end reflection was improved from approximately 200 m

to 50 m. This is the same spatial resolution as that reported in our previous study under the same

frequency-sweep rate y [6], where we narrowed the time-gate width to less than 0.5 ms but

sacrificed the measurement time. The ratio of the spatial resolutions to the fiber length were 1 m

/I km = 10"-' and 50 m /30 km = 1.7 x 10 ^ which are the best results to date for C-OFDR.

4. Conclusion

We demonstrated C-OFDR for long optical fibers using externally modulated lightwaves.

By suppressing the frequency-sweep nonlinearity with an auxiliary feedback-control circuit, we

achieved spatial resolutions of 1 m and 50 m for the far-end reflection from 1 km and 30 km

optical fibers, respectively, without substantial narrowing of the time-gate width in the spectrum

analysis.

References

[1] U. Glombitza and E. Brinkmeyer, "Coherent frequency-domain reflectometry for

characterization of single-mode integrated-optical waveguides", J. Lightwave TechnoL, vol. 11,

pp. 1377-1384, 1993

[2] R. Passy, N. Gisin, J. P. von der Weid, and H. H. Gilgen, "Experimental and theoretical

investigations of coherent OFDR with semiconductor laser sources", J. Lightwave TechnoL, vol.

12, pp. 1622-1630, 1994

[3] W. V. Sorin, D. K. Donald, S. A. Newton, and M. Nazarathy, "Coherent FMCW reflectometry

using a temperature tuned Nd:YAG ring laser", IEEE Photon. TechnoL Lett., vol. 2, pp. 902-904,

1990

[4] C. W. Lee, E. T. Peng, and C. B. Su, "Optical homodyne frequency domain reflectometer

using an external cavity semiconductor laser", IEEE Photon. TechnoL Lett., vol. 7, pp. 664-666,

1995

[5] K. Tsuji, K. Shimizu, T. Horiguchi, and Y. Koyamada, "Coherent optical frequency domain

reflectometry for a long single-mode optical fiber using a coherent lightwave source and an

external phase modulator", IEEE Photon. TechnoL Lett., vol. 7, pp. 804-806, 1995

[6] K. Tsuji, K. Shimizu, T. Horiguchi, and Y. Koyamada, "Coherent OFDR using phasedecorrelated lightwaves for characterizing optical fiber transmission lines", Ilth International

Conference on Optical Fiber Sensors (OFS-II ), pp. 804-806, 1996

205
Z\v=3kHz

Er-Yb-glass EO

Laser

Modulation

Frequency

A 7=0.1 THz/s

2 GHz

Delay Fiber

500 m

through

Synthesized

Sweeper

Mixer

FD

/X/

/\/ 0R2

^Time

20 msec

AMP LPF

Figure 1 . Experimental setup.

BPF

RF Signal Source

78.5 79.0 79.5 80.0 80.5 81.0 81.5 Frequency (MHz)

Figure 2. Spectrum of the beat signal from optical receiver 2 (0R2).

Time (2 ms/div.)

Figure 3. Frequency fluctuation of the reference

signals . Solid and dotted lines correspond to the traces with and without the frequency-sweep

control, respectively.

1 1

(a)-

W91

(b)i

1 . 81.068 81.072 81.076 81.080 81.084

Frequency (MHz)

Figure 4. OFDR waveforms for the vicinity of the

far end of the 1 -km-test fiber (a) without, and (b) with the frequency-sweep control.

1 ! r

110.8 110.9 111.0 111.1 111.2 Frequency (MHz)

Figure 5. OFDR waveforms for the far end of the 30-km-test fiber (a) without, and (b) with the frequency-sweep control.

206
A NEW WroE SCAN RANGE HIGH RESOLUTION

OPTICAL LOW COHERENCE REFLECTOMETER

Mauro Bottanelli

SIRTI S.p.A., Cables and Optical Technologies

Via Manzoni 44, 20095 Cusano Milanino (MI), ITALY

Abstract

A transportable and easy to use optical low coherence reflectometer has been development for fault detection and optical component characterization. Up to 40 centimeters scan range, better than 80 |am spatial resolution and more than 100 dB dynamic range have been achieved by means of a new measurement

configuration, without any polarization control elements or signal filtering technique.

Introduction

Optical Low Coherence Reflectometry (OLCR) has been already pointed out in the past as a suitable technique for high resolution fault detection and optical component characterization. Recently, the availability

of erbium doped optical amplifiers as light soiu-ces and the growing interest in the development and testing of

high technology optical components (couplers, WDMs, filters, attenuators, isolators...) have emphasized OLCR

potentiality with respect to traditional measurement method. Commercially available Optical Domain

Reflectometers (OTDR), for instance, are limited both in resolution and in dynamic range when a centimetric or a sub-centimetric fault location is needed, or when a very low back-reflection component has to be analysed.

On the contrary, OLCR systems offer suitable features for those goals.

This paper describes the design, the development and the first characterization of a new transportable

and easy to use OLCR set-up, which has been conceived both for field fault detection and for optical component

analysis and quality control applications. The new OLCR set-up has been developed by SIRTI R&D

laboratories in cooperation with the Polytechnic of Milan during the last year.

Theory and background

Fig. 1 reports the basic scheme of a generic OLCR set-up. The continuous wave low coherence radiation

from the light source is parted into a measiu'ement and a reference beam by means of a beam splitter. After

reflections, respectively inside the Device Under Test (DUT) and from the variable length reference path

mirror, the beams are combined by the splitter, giving interference patterns on the detector. The set-up

therefore reproduces a Michelson interferometer, but, due to the source low coherence time, interference can

rise only for very limited difference between the measurement and the reference path, that is: discrimination

among different back-scattering points inside the DUT can be obtained with a very high spatial resolution,

simply moving the reference mirror by means of a micropositioning stage. The larger the source optical power,

the higher the OLCR set-up dynamic range; the narrower the source autocorrelation fimction, the higher the OLCR set-up resolution. In the past, only Light Emitting Diodes (LEDs) or Super-Luminescent Diodes (SLDs)

constituted smtable sources for OLCR set-ups, but better performances can be achieved nowadays by means of

the Amplified Spontaneous Emission (ASE) of an erbiiun doped Optical Fiber Amplifier (OFA).

Let us now review the main featiu-es of OLCR systems:

• Spatial resolution : the shortest spatial distance between two contiguous scattering points that can be

distinguished by the system. It can be proved to be:

(1) Az = 2«AA

where X and Ak are respectively the wavelength and the spectial width of the source, and n is the material

refraction index of the DUT.

• Sensitivity: the lowest back-scattered signal that can be detected by the system. It can be proved to be:

207
a) J\im ~ (^'^^^ limited case)

Ifdc

(2)

b) = —- (excess noise limited case)

2y • ov

being B the detection band of the optical signal, q the elementary charge, y the DUT insertion loss, I^j^ the photodetector direct current offset and 5v a sort of equivalent source band width defined by the source power spectrum G(v) as:

+00 n2

J G{v)dv

+00 \G^{v)dv

In the first case, the system sensitivity is mainly limited by the photodetector characteristics; in the second,

it is limited by the light source intensity fluctuations.

• Blind space : the spatial width of some areas inside the DUT where the system can not perform a good

measurement, due to the presence of minor lobes in the source autocorrelation fimction. Those areas are

typically adjoining intense light back-scattering centers in the DUT. Their spatial extension can range

from 20 mm (multimode LD source) to 5 mm (LED, SLD source) or to 0.5 mm (ASE source) considering

the source FW60dB (Full Width at 60 dB fi-om the top of the autocorrelation fimction).

• Scan ranse : the largest continuous spatial range that can be scaimed by the system, typically about 20

centimeters more or less, depending on the system design and purpose. A new OLCR configuratioii

Fig. 2 sketches the main features of our new OLCR configuration. The light source is the amplified

spontaneous emission of an erbium doped optical amplifier, having 6 dBm optical output power at 1530 nm,

about 8 mm FWHM spectral width, a 1 ps FWHM autocorrelation fimction width (see figure 3) and a very low

polarization degree. Light is launched into a traditional SMR type optical fiber without any particular kind of

polarization control, and is parted into measurement and reference beams by means of a 50:50 fiision coupler.

A sawtooth ramped piezoelectric cylinder (PZT) as phase modulator (PM) has been inserted on the OLCR

reference arm in order to improve the sensitivity by means of a lock-in amplifier detection. A particular

micropositioning motorized stage (up to 0.2 micron step resolution) can adapt the reference path to the measiu'ement one, in order to scan up to 40 cm DUT length. Other two 50:50 fiision couplers and two InGaAs

PIN detectors have been inserted into the configuration in order to implement the so-called "balanced

detection" technique, that is a subtraction between photocurrents deleting the source intensity fluctuation excess

noise. In ideal conditions (coupler splitting ratio k=0.5, equal detector responsivities, identical detection pigtail

paths) the system can be considered shot noise limited only, with even up to a 40 dB sensitivity improvement

with respect to the single PIN detection. The opto-mechanical set-up is completely controlled by a personal

computer by means of a C language developed program.

A comparison among our new set-up and the most effective traditional OLCR configurations [1, 2, 3]

(fig. 4/a, /b and /c) is simimarized in the following points:

• in the case of a perfectly balanced detection, /b configuration sensitivity can be proved to be theoretically the highest; our configuration and the c/ one have a theoretical sensitivity 3 dB worse than for /b. /a

configuration can not be considered balanced in any case (it is the worst one). However the real sensitivity also depends on the PZT position: due to macrobending losses, it can be proved that the PZT placed on the

208
measurement ann can make sensitivity worse, while, as in our configuration, the PCT placed on the

reference arm does not influence the sensitivity;

• our and /a configm-ations have the highest scan range, because they double the moving stage mechanical

scan range. Moreover, as far as we know, a 40 cm scan range has never been reported in the Uterature till now;

• our and /a configurations delete the influence of possible light reflections at the fiber-air interfaces, because

in these two configurations imwanted reflected light and the reference light travel along the same physical

paths.

New OLCR set-up characterization and Hrst experimental results Table 1 reports out OLCR set-up main features, as measured according to several trial experimental

results. It is interesting to point out the fact that these such remarkable performances are however still under

the maximum configm-ation theoretical potentialities (the maximum theoretical sensitivity level, for instance,

has been caloilated to be 142 dB), and that they have been achieved without any kind of particular optical

component selection, polarization control systems or hardware end software filtering technique. Those present

featiu"es can be considered to be sufficient for every practical measurement purposes, a part from fiber Rayleigh

scattering level evaluation, which would require an higher sensitivity and, in general, a very long measurement

time.

Sensitivity 107 dB

Repeatability 0,86 dB

Reproducibility 0,89 dB

Spatial resolution (FWHM) 75 |j,m

FW20dB = 260 ^m

Blind space FW40dB = 550 ^m

FW60dB = 1,04 mm

Scan range in the air 60 cm

Scan range in the fiber 40 cm

Measurement mean time 5 minutes

(for a 2 cm long DUT) (at the finest scan step)

Table 1 - New OLCR set-up mean features

By means of an experimental report, fig. 5 illustiates the capability of the new OLCR system to be used

as fault detector. Other measurement examples will be shown during the paper presentation, especially

concerning component characterization.

Conclusions

A new OLCR set-up for fault detection and high resolution component characterization has been

developed and described. Up to 40 centimeters scan range, better than 80 jim spatial resolution and more than

100 dB dynamic range have been achieved.

Acknowledgements

This work has been developed by the author and by Adolfo Ottonello during a one year stage in SIRTI's

laboratories. We would like to thank Mario Martinelli (Polytechnic of Milan) for his invaluable help in

theoretical question deepening.

References

[1] H.H. Gilgen, G. Bodmer, C. Zimmer, "Optical coherence domain rejlectometry as a test method for

integrated optics devices", 2nd Optical Fibre Measm-ements Conference, Turin, p. 143-146, 1993.

[2] W.V. Sorin, D.M. Baney, "Measurement of Rayleigh backscattering at 1.55 /i/w with 32 ^im spatial

resolution", Photonics Technology Letters, vol. 4, p. 374-376, 1992.

[3] K. Takada, T.Kitagawa, M. Shinizu, M. Horiguchi, "High-sensitivity low coherence reflectometer using

erbium-doped superfluorescentfibre source and erbium-doped power amplifier", Electronics letters, vol. 29, p. 365-367, 1993.

209
Fig. 2 - Sketch of the new OLCR set-up

rtirror displacensnt Cprt]

Fig. 3 - ASE source autocorrelation function Fig. 4 - Most eflfertive traditional OLCR set-ups

0

-20

I

I

-60

I -80

-100

-120

noise floor

connection fiber break •

CD CO to 00 (O oo CM CO in h~ CD CO 00 CO CD CO in 00 CD o CO in CD 00 CD o CD in o in o oo" o m o CO CD_ CD to CVJ in cd' o CO in to" oo" csT in t^' oo" O co' CM CO CO CO CO CO CO m m in

position [mm]

Fig. 5 - Measurement report example: fiber break identification

210
Interferometric noise in fiber transmission systems incorporating birefringent substances

Michael A. Bukhshtab

Coming Incorporated, Opto-Electronic Components, Coming, NY 14831

The impact of interferometric noise terms on the bit-error-rate floors of a transmission system with

a DFB laser degrades its performance due to the conversion of laser phase noise into signal dependent

intensity noise [1]. Measurement systems evaluating optical amplifiers, dispersion compensation fibers

and WDM elements often contain birefringent components. Transmission of a coherent and polarized

laser beam through a birefringent substance between potential refractive index discontinuities creates a

complex interference pattern contributing to auxiliary intensity noise of the measurement system itself.

This paper presents a computation of interference noise in a distributed birefringent substance by

summing interference patterns created by components with equal and different numbers of reflections

along each birefringent axis, and then within any selected polarization direction. The optimum phase

settings for measurements under low auxiliary noise conditions are identified by computing the phase

retardance of a total system that is required to establish its performance as a noise filter. This approach

assumes the similar frequency dependence and probability distribution of interferometric noise as in [1]

The multiple-path interference in a birefringent substance can be represented as two processes for

waves with equal and different numbers of passes through it: the comparative and the cross-reference

events [2]. OrJy light modes that contribute to the comparative interference have identical amplitudes.

Presuming that the products of both surface reflections of this substance are (poipo2)^, (Peipe2 )^ «1 only

the interference of waves separated by at least two dual-reflection circles is not counted. Such a secondorder approximation, computing interference among all components diverted by any dual pass, gives:

(1)

,r (2m-l)5

2

+ p:pr'sin^fm5„+|j-hp.

211
here 5 = (27i:/X,)(ne-no)h; h is thickness of the substance; po, pe are its reflectances along the ordinary and

extraordinary axes; Om =(2m + l)(5e - 5o); 5o = (27c/?i)(hno), 8e = (27c/?i)(hne); m = 1, 2, 3... is the integer.

The first term in equation (1) accounts for the comparative interference patterns along the ordinary

and extraordinary axes of the birefringent substance. The second term includes some infinite number

of components with different amplitudes, which participate in the cross-reference interference between

the ordinary and extraordinary directions while having mo and m^ reflections, respectively. Let us omit

interference of waves separated by more than one dual-reflection circle inside the retarder. This means

that squared surface reflection factors: (poiPo2)^, (peipe2 )^ are negligible compared to uruty. For identical

front and back surface reflections this is equal to (po^)^ « 1/ (Pe^)-^ « 1- Equation (1) thus transforms to:

-\2

V0-Po)'+4p„ sin'5„

cos(pcos((p -7)+

^(l-pj'-h4p,sin'5,

sin(psin((p -

y)

(^] f

+ sin' + P.sin'

.2, V

sin

+ p'p. sin'

6„ + —

2

+ 9e sin^ 5 +

5^ 1 +

26 +

5^

+ P«p'sin' 26 +

6^

(2)

sin 2(p sin [2(9 -y)].

Equation (2) still counts many multiple reflections. When boundary reflectances are low, the number of

accepting terms could be reduced by the approximation of poiPo2 « 1, Peipe2 « 1 or po^ « 1, Pe^ « 1:

1-p. 1-p.

-i2

A/(l-P.)'+4p<,sin'6,

cos(pcos((p -y)-lVO-PJ'+^P. sin' 6,

-sinq)sin(cp -

y)

(3)

-/oT.Tjsin^ ^-1

2y

^36^

6„-h2+ p, sin^

2

sin2(psin[2((p-y)]

Assuming a limited coherence length, when wave components with phase shifts 60, 6e do not interfere:

^ ^cos(pcos((p -y) +

^| | sin(psin((p -y)

1 + P.

-/o(l-pJ(l-p.)|sin^

r UJ

+ 9o9e sin'

V

r36^

K ^ J

1 + P.

Uin2(psin[2((p -y)]

(4)

Fig. 1 depicts a quarter-wave retarder for X,=1555 nm with spectral attributes similar to quartz whose

optical axis is oriented at 30° to parallel input and output polarization directions, so y = 0°; (p = 30°. These

212
computations are based on a presumption of six percent reflections for polished fiber endfaces as in [1].

Effects are computed by equations (2) - (4) (curves 1-3) versus an ideal case of no reflections (curve 4).

Although the character of each curve is similar, the equations compute different intensity noise, seen as

oscillations of mean intensities. Since at low reflectances equations (2), (3) give similar results (curves 1,

2), there is no need to consider all multiple reflections. However, the reduced amount of determined by

equation (4) terms (curve 3) predicts an incorrect light transmission. It also is seen that interference noise

is elevated at high and low intensities of transmitted light, related to multiples of 7i/2 phase retardance.

It is different near the 0.5 intensity regions matching ideally optimal 8 opt = tc/4. Counting all reflections

the actual 5opt point changes, since dual-path phase shifts for each optical axis are not multiples of 7c/4.

Fig. 1 . Transmission of the quarter-wave retarder for 1 555 nm between two air gaps:

1 - all effects are counted; 2 - only first two dual-reflection cycles are included; 3 - only

two-beam interference is considered, 4 - ideal case assuming no reflections

•Seriesi Series2

-Series3 Series4

Wavelength, microns

Let us thus define that optimal phase shift 5opt, which is needed to obtain the minimum influence

of spectral changes of incident light, and correspondingly the lowest interference noise in transmission:

^[J.xjuuilh) _ _VC^JsinS +3p„p, sin36 + p„ sin(25„ sin(25, +5)-

dd 2 [+pJp^sin(46„-+-5)+p„p;sin(45,+5)

>sin2(psin[2((p -y)] (5)

{sin 5 +3p„p, sin38 -i-p„sin(25„ -i-5)-i- p^ sin(25, -F6)}sin 29 sin[2((p -y )} (6)

(7)

95 2

_v 1—/ ^ _ V— —rw I g ^ 3p^p^ ^jj^ 3g 2(p sin 2((p -

y ) L do 2

The potential points of the lowest interference changes of the observing intensity, and thus minimal

interference noise, are corresponding to zero of the second derivatives. These second derivatives are: 213
2

COS 5 + 9p„ COS 35 + p,, cos(25„ + S ) + p^ cos(25^ + 6 )

+ p,'p, cos(45„ +5) + p,p,' cos(45, +5)

sin2(psin[2((p-y)].(8)

= -^{cos5 +9p„p, cos38 + p„ cos(25^ +6)+ p^ cos(25, +5)}sin2(psin[2((p -y)J (9)

^^{hx.ssUuncoherlh) (l - PJ(l " pJ {cos5 + 9p„p^ cos35}sin2(psin[2((p -y)} (10)

2

The first and second derivatives of transmission given by equations (5) - (10) are depicted by Fig. 2, 3.

0.44

0.436

(D

SS 0.424

Fig. 2. Intensity changes in transmission at the quarter-wave retardance at 1555 nm existing

between refractive index discontinuities: curves 1 - 3 are computed to equations (5) - (7)

0.42 +

0.416

A A 3 ^

Region of the low

1,2 intensity changes

1 i 1 .1 J - - .1

lO

Wavelength, microns

Fig. 3. Finding the optimal wavelength that matches the optimal phase retardance and

the lowest excessive noise: curves 1 - 3 are computed by equations (8) - (10)

(0 SI

a.c

o

o

0)

0.06

0.04

0.02

0

-0.06

" 1,2

3 The lowest ^

noise point

1

L ) 00

1

If) —1

CM in ^F?«iaKTMsas,

1

CD

1

Cvl u>

^: ^

in in in ^ CO Cl>

•Seriesi Series2

- Series3

Seriesi

Series2

- SeriesS

Wavelength, microns

The proposed reduction of birefringence-induced multiple-path interference noise terms can expand

design of transmission and measurement systems by incorporating fibers with the optimal birefringence.

Reference:

[1] J. L. Gimlett, N. K. Cheung, Effects of phase-to-intensity noise conversion by multiple reflections on

gigabit-per-second DFB laser transmission systems,]. Lightwave Technol., 1989, Vol. 7, No. 6, pp. 888-895

[2] M. Bukhshtab, The influence of surface reflections on computation and measurement of retardance,

Meas. Sci. Technol., 1995, Vol. 6, No. 7, pp. 910 - 920

214
AUTHOR INDEX

Anderson, Duwayne ... 191

Antos, A.J 61

Artiglia, M 31

Arvidsson, Bertil 139

Aso, Osamu 147,151

Auguste, J.L 187

Barlow, A.J 151

Bartelt, H 27

Berkey, G.E 61

Blondel,M 195

Blondy, J.M 175, 187

Boness, R 75

Bottanelli, Mauro 207

Brising, Kim 139

Bukhshtab, Michael A. 211

Chia, Shin-Lo 57

Clapeau,M 187

Clayton, J.B 105

da Silva, Valeria L 61

Daniaux, D 195

Del Core, Andrea 67

Dell, J.M 35

DiBin, P 183

Dultz, W 71

Dupre, J 1 Elamari, A 163

Ertmer, A 27

Eiselt,M 115

Facq,P 175, 183, 187

Faugeras, P 183

Fellegara, A 67, 199

Franzen, D.L 121

Fujishima, H 151

Fuss, M 13

George, G 89

Gisin, B 131

Gisin,N. 131,143,151,163

Hackert, Michael J 79

Headley, Clifford 105

Heens, B 31

Heffner,B.L 135

Herchenroder, G. ... 71, 93

Horiguchi, T 203

Humphreys, D.A 97

Ichimura, M 151

Ishigure, Takaaki 171

Jopson, R.M 115

Judy,A.F 151

Julliard, K 143

Kashyap, Raman 7

Keenan, Darryl A 49

Kim,K.S 151

Kirchhof,J 75

Klein, K.-F 75

Koike, Yashuhiro 171

Koyamada, Y 203

Krusius,J.P 89

Larsen, C 31

Li,Xaioyu 49, 101

Liu,Y 61

Martinelli, Mario 67

Mechels, S.E 121

Megret, P 195

Melloni, Andrea 67

Mickelsson, Hans 139

Mitachi, S 53

Moeyaert, V 195

M6rl,K 93

Muller, H.-R 27

Nagase, R 53

Namihira, Yoshinori 151, 179

Newhouse, M.A 61

Nihei, Eisuke 171

Nikles, Marc Ill

Noguchi, M 151

Oberson,Ph 143

Ogoshi, Haruki 147

Ohashi,M 151

Ozeki,T 151

Pagnoux,D. . 175, 183, 187

Passy, R 131, 143

Pemy,B 151,163

Potenza, Marcello 23

Rasmussen, Jens 167

Reed, W.A 105

Reichel,V 27

Reith, L.A 39

Robert, Philippe A. ... Ill

Rothhardt, M 27

Roy, P 175

Salina, Pascal Ill

Saravanos, Costas 45

Schlager, J.B 121

SchoU, Bernhard 167

Schuh, R.E 159

Scott, Thomas R. . . 49, 101

Shimizu, K 203

Siddiqui, A.S 159

Siliquini, J.F 35

Simos, C 183

Smith, David 83

Soccolich, C 31

Suetsugu, Y 151

Sutherland, J 89

Suzuki, S 151

Sawada, M 151

Tambosso, Tiziana 23

Thevenaz, Luc Ill

Townley-Smith, Paul . . 45

Tsuji, K 203

Unger, S 75

Vayshenker, Igor . . 49, 101

Vobian, J. , 71,75,93

Von der Weid, J.P. 131,143

Willems, F.W 17,31

Williams, P.A. 125, 151, 155

Wittl,F 71,75

Wuilmart, L 195

Xu,M.G 35

Yang, Shao 101

Zainul, M.Ajaaz 79

Zbinden, H 163

Zimmer, Ch 1 63

215 •U.S. GOVERNMENT PRINTING OFFICE: 1996-404-525/67902

Technical Publications

Periodical

Journal of Research of the National Institute of Standards and Technology—Reports NIST research

and development in those disciplines of the physical and engineering sciences in which the Institute is active. These include physics, chemistry, engineering, mathematics, and computer sciences. Papers

cover a broad range of subjects, with major emphasis on measurement methodology and the basic

technology underlying standardization. Also included from time to time are survey articles on topics

closely related to the Institute's technical and scientific programs. Issued six times a year.

Nonperiodicals

Monographs—Major contributions to the technical literature on various subjects related to the Institute's scientific and technical activities. Handbooks—Recommended codes of engineering and industrial practice (including safety codes)

developed in cooperation with interested industries, professional organizations, and regulatory bodies.

Special Publications—Include proceedings of conferences sponsored by NIST, NIST annual reports, and

other special publications appropriate to this grouping such as wall charts, pocket cards, and

bibliographies.

Applied Mathematics Series—Mathematical tables, manuals, and studies of special interest to physicists,

engineers, chemists, biologists, mathematicians, computer programmers, and others engaged in scientific and technical work.

National Standard Reference Data Series—Provides quantitative data on the physical and chemical

properties of materials, compiled from the world's literature and critically evaluated. Developed under a worldwide program coordinated by NIST under the authority of the National Standard Data Act (Public

Law 90-396). NOTE: The Journal of Physical and Chemical Reference Data (JPCRD) is published bi- monthly for NIST by the American Chemical Society (ACS) and the American Institute of Physics (AIP).

Subscriptions, reprints, and supplements are available from ACS, 1155 Sixteenth St., NW, Washington,

DC 20056.

Building Science Series—Disseminates technical information developed at the Institute on building

materials, components, systems, and whole structures. The series presents research results, test methods, and performance criteria related to the structural and environmental functions and the durability

and safety characteristics of building elements and systems.

Technical Notes—Studies or reports which are complete in themselves but restrictive in their treatment of a subject. Analogous to monographs but not so comprehensive in scope or definitive in treatment of

the subject area. Often serve as a vehicle for final reports of work performed at NIST under the sponsorship of other government agencies.

Voluntary Product Standards—Developed under procedures published by the Department of Commerce

in Part 10, Title 15, of the Code of Federal Regulations. The standards establish nationally recognized

requirements for products, and provide all concerned interests with a basis for common understanding of

the characteristics of the products. NIST administers this program in support of the efforts of privatesector standardizing organizations.

Consumer Information Series—Practical information, based on NIST research and experience, covering

areas of interest to the consumer. Easily understandable language and illustrations provide useful background knowledge for shopping in today's technological marketplace.

Order the above NIST publications from: Superintendent of Documents, Government Printing Office,

Washington, DC 20402.

Order the following NISTpublications—FIPS and NISTIRs—from the National Technical Information

Service, Springfield, VA 22161.

Federal Information Processing Standards Publications (FIPS PUB)—Publications in this series collectively constitute the Federal Information Processing Standards Register. The Register serves as the

official source of information in the Federal Government regarding standards issued by NIST pursuant to the Federal Property and Administrative Services Act of 1949 as amended, Public Law 89-306 (79 Stat. 1127), and as implemented by Executive Order 11717 (38 FR 12315, dated May 11, 1973) and Part 6 of

Title 15 CFR (Code of Federal Regulations).

NIST Interagency Reports (NISTIR)—A special series of interim or final reports on work performed by

NIST for outside sponsors (both government and non-government). In general, initial distribution is handled by the sponsor; public distribution is by the National Technical Information Service, Springfield,

VA 22161, in paper copy or microfiche form.